Advances in Cryogenic Engineering VOLUME 211 A Continuation Order Plan is available for this series. A continuation order will bring delivery of each new volume immedlately upon publicatlon. Volumes are billed only upon actual shipment. For further Information please contact the publlsher. A Cryogenic Engineering Conference Publication Advances in Cryogenic Engineering VOLUME 25 Edlted by K. D. Tlmmerhaus Engineering Research Center University of Colorado Boulder, Colorado and H. A. Snyder Department of Aerospace Engineering Seiences University of Colorado Boulder, Colorado SPRINGER SCIENCE+BUSINESS MEDIA, LLC The Library of Congress cataloged the first volume of this title as follows: Advances in cryogenic engineering. v. 1New York, Cryogenic Engineering Conference; distributed by Plenum Press, 1960v. illus., diagrs. 26 cm. are reprints of the Proceedings of the Cryogenic Engineering Vols. 1Conference, 1954K. O. Timmerhaus. Editor: 19601. Low temperature engineering-Congresses. ed. 11. Cryogenic Engineering Conference. 1. Timmerhaus, K. 0., 660.29368 TP490.A3 57-35598 Proceedings of the 1979 Cryogenic Engineering Conference, held at the University of Wisconsin, Madison, Wisconsin, August 21-24, 1979. Library of Congress Catalog Card Number 57-33598 ISBN 978-1-4613-9858-5 ISBN 978-1-4613-9856-1 (eBook) DOI 10.1007/978-1-4613-9856-1 © 1980 Springer Science+Business Media New York Originally published by Plenum Press, New York in 1980 Softcover re print of the hardcover 1st edition 1980 Ali rights reserved No part of this book may be reproduced, stored in a retrieval system, or transmittild, in any form or by any means, electronic, mechanical, photocopying, microfilming, recording, or otherwise, without written permission from the Publisher CONTENTS Foreward ........................................................ . Russen B. Scott Memorial Award ................................... . Samuel C. Collins Award .......................................... . 1979 Cryogenic Engineering Conference Board ....................... . Awards Committees ............................................... . Acknowledgments ................................................ . xii xiii xiv XV XV XVI Superconductivity Applications-MHD and Fusion Magnets A-1 A-2 A-3 A-4 A-5 A-6 Superconducting MHD Magnet Engineering Program, P. G. MARSTON, A. M. DAwsoN, D. B. MoNTGOMERY, and J. E. C. WILLIAMS, Francis Bitter National Magnet Labaratory . . . . . . . . . . . Impact of High-Current Operation on the Cost of Superconducting Magnet Systems for Large-Scale MHD Applications, R. J. THOME, R. D. PILLSBURY, H. R. SEGAL, and B. 0. PEDERSON, Magnetic Corporation of America . . . . . . . . . . . . . Final Design of a Superconducting MHD Magnet for the CoalFired Flow Facility at the University of Tennessee Space Institute, S.-T. WANG, L. R. TuRNER, L. GENENS, W. PELCZARSKI, J. HoFFMAN, J. GoNCZY, H. Luowm, R. C. NIEMANN, K. F. MATAYA, and E. KRAFT, Argonne National Laboratory, and W. YouNG, University of Wisconsin . . . . . . . . . . . . Cryogenic Aspects of the UTSI-CFFF Superconducting Dipole Magnet for MHD Research, R. C. NIEMANN, S.-T. WANG, J. W. DAWSON, L. GENENS, R. P. SMITH, L. R. TURNER, J. D. GoNCZY, J. HOFFMAN, and K. F. MATAYA, Argonne National Laboratory, P. SMELSER, Independent Consultant, and P. C. VANDER AREND and S. STOY, Cryogenic Consultants, Inc. . . . . . . Safety Analysis of the UTSI-CFFF Superconducting Magnet, L. R. TuRNER, S.-T. WANG, and R. P. SMITH, Argonne National Laboratory, P. C. VANDER AREND, Cryogenic Consultants, Inc., and Y.-H. Hsu, General Atomic Company . . . . . . . . . . . . . . . . Engineering Aspects of Cryogenic Laser-Fusion Targets, D. L. MusiNSKI, T. M. HENDERSON, R. J. SIMMS, and T. R. PATTINSON, KMS Fusion, Inc., and R. B. JAcoas, R. B. Jacobs Associates, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 12 19 30 39 49 Superconductivity Applications-Energy Transfer and Storage B-1 Energy Transfer in a System of Superconductive Magnets, M. MASUDA, T. SHINTOMI, and K. AsAJI, National Labaratory for High Energy Physics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . V 61 vi B-2 B-3 B--4 B-5 B-6 B-7 B-8 Contents Thermal Cycle Tests of a Modeted Superconducting Transmission Line, C. F. SINDT and P. R. LuoTKE, NBS Thermophysical Properties Division . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Design of a 400-kJ Pulsed Energy Storage Coil, S. K. SINGH, C. J. HEYNE, D. T. HACKWORTH, M. A. JANOCKO, P. W. EcKELS, and J. H. MURPHY, Westinghouse Electric Corporation . . . . . . . . . . . . . . Operating Characteristics of a 1.5-MJ Pulsed Superconducting Coil, S. H. KIM, S.-T. WANG, and M. LIEBERG, Argonne National Labaratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-MJ Magnet for Superconductive Energy Storage, T. SHINTOMI, M. MAsUDA, H. SATO, and K. AsAn, National Labaratory for High Energy Physics . . . . . . . . . . . . . . . . . . . . . . . . . . . Conceptual Design of a 20-MJ Superconducting Forced-Cooled Ohmic-Heating Coil, S. K. SINGH, J. H. MURPHY, M. A. JANOCKO, H. E. HALLER, D. C. LITz, and P. W. ·EcKELS, Westinghouse Electric Corporation, and J. D. RoGERS and P. TRULLEN, Los Alamos Scientific Labaratory . . . . . . . . . . . . . . . . . . . . Shape Optimization Study for a Three-Tunnel Superconductive .Iinergy Storage Magnet, M. N. EL-DERINI, University of Petroleum and Minerals, and R. W. BooM, University of Wisconsin . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20-kA Power Supply for Large Superconductive Coils, M. MASUDA, T. SHINTOMI, and K. AsAn, National Labaratory for High Energy Physics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69 81 90 98 105 114 120 Superconductivity Applications-Rotating Machinery C-1 C-2 C-3 C--4 Superconducting Generator Design for Airborne Applications, B. B. GAMBLE and T. A. KEIM, General Electric Company......... Superconducting Field Winding for a 10-MVA Generator, K. A. TEPPER, J. V. MINERVINI, and J. L. SMITH, Jr., Massachusetts Institute of Technology............. .......................... Conductive Armature Shielding Design Concepts for Slow-Speed Superconducting Generators in the 40- to 400-MVA Range, S. KuzNETSOV, Imperial College of Science and Technology......... Optimization of Superconducting Cryoturbogenerator FieldWinding Parameters, B. I. VERKIN, I. S. ZHITOMIRSKII, and R. V. GAVRILOV, Academy of Seiences of the Ukrainian SSR . . . . . . 127 137 145 156 Superconductivity Applications-Magnet Technology D-1 D-2 D-3 Superconducting Compensating Solenoids for the CELLO Detector Experiment at PETRA, W. BARTH, N. FESSLER-WILHELMI, W. LEHMANN, and P. TUROWSKI, Kernforschungszentrum Karlsruhe Construction and Test of the CELLO Thin-Wall Solenoid, H. DESPORTES, J. LE BARS, and G. MAYAUX, Centre d'Etudes Nuc/eaires de Sac/ay . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Quenches in the Superconducting Magnet CELLO, W. V. HASSENZAHL, Los Alamos Scientific Labaratory . . . . . . . . . . . . . . . . 163 175 185 Contents D-4 D-5 D-6 D-7 D-8 Construction and Testing of the Two-Meter-Diameter TPC Thin Superconducting Solenoid, M. A. GREEN, P. H. EBERHARD, R. R. Ross, and J. D. TAYLOR, Lawrence Berkeley Laboratory . . . . . Superconducting Magnet System for the Spirit Cosmic Ray Space Telescope, M. A. GREEN and J. M. DEÜLIVARES, Lawrence Berkeley Laboratory, and G. TARLE, P. B. PRICE, and E. K. SHIRK, University of Ca/ifornia, Berkeley . . . . . . . . . . . . . . . . . . . . . . A Maintainable Superconducting Magnet System for Tokamak Fusion Reactors, S. Y. HsiEH, G. DANBY, J. R. PowELL, P. BEZLER, D. GARDNER, and C. LAVERICK, Brookhaven National Laboratory, and M. FINKELMAN, T. BROWN, J. BuNDY, T. BALDERES, I. ZATZ, R. VERZERA, and R. HERBERMAN, Grumman Aerospace Corporation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Prototype Low-Current Superconducting Quadrupole Magnet for Fermilab's High-Intensity Laboratory, W. CRADDOCK, R. W. FAST, P. GARBINCIUS, and L. MAPALO, Fermi National Accelerator Laboratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Superconducting Magnets of the Biomedical Facility at SIN, J. ZELLWEGER, G. VECSEY, and I. HoRVATH, SIN Swiss Institute for Nuclear Research . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vü 194 200 207 222 232 Superconductivity Applications-Cryogenic Techniques E-1 E-2 E-3 E-4 A Novel Thermometer Sensor for the mK Region Using the Proximity Effect, H. NAGANO, Y. ÜDA, and G. Fum, Tokyo University . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A Superconducting RF Notch Filter, C. S. PANG, C. M. FALCO, R. T. KAMPWIRTH, and I. K. ScHULLER, Argonne National Laboratory, and J. J. HuoAK and T. A. ANASTASIO, Department of Defense . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Experimental Evaluation of a 1-Meter-Scale D-Shaped Test Coil Fabricated from a 23-Meter Length of Internally Cooled, Cabled Superconductor, M. 0. HoENIG, A. G. MONTGOMERY, and S. J. W ALDMAN, Francis Bitter National Magnet Laboratory . . . . . . . . . . Performance of Gas-Filled Thermal Switches, J. YAMAMOTO, Osaka University. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 239 244 251 261 Cooling Superconducting Systems F-1 F-2 F-3 F-4 Transient Cooling of a Faultworthy Superconducting Electric Generator, J. A. ScHWOERER and J. L. SMITH, Jr., Massachusetts Institute of Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Experimental Simulation of a Cryogenic System for a Large Superconducting Rotor, L. SOBEL, J. L. SMITH, Jr., and F. RuMORE, Massachusetts Institute of Technology . . . . . . . . . . . . . . . . Rotor Cooling System for a 10-MVA Superconducting Generator, M. T. BROWN, M. E. CRAWFORD, and J. L. SMITH, JR., Massachusetts Institute of Technology . . . . . . . . . . . . . . . . . . . . . . Safety Leads, M. KucHNIR and T. H. NICOL, Fermi National Accelerator Laboratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 266 275 285 294 Coateats F-5 Magnet Leads for the First-Cell, D. P. ßROWN and W. J. ScHNEIDER, Brookhaven National Labaratory . . . . . . . . . . . . . . . . . . F-6 Thermal Control for the MFfF Magnet, J. H. VANSANT and R. M. Russ, Lawrence Livermore Labaratory . . . . . . . . . . . . . . . . . . . . . . F-7 Forced-Circulation Cooling System for the Argonne Superconducting Heavy-Ion Linac, J. M. NIXON and L. M. BoLLINGER, Argonne National Labaratory . . . . . . . . . . . . . . . . . . . . . F-8 Energy Doubler Satellite Refrigerator Magnet Cooling System, C. RODE, P. ßRINDZA, and D. RICHIED, Fermi National Accelerator Laboratory, and S. STOY, Cryogenic Consultants, Inc. . . . . . . . . . . . . F-9 Cryogenic Support System for Airborne Superconducting Generators, P. J. KERNEY and P. A. LESSARD, CTI-Cryogenics . . F-10 Economics of Cryogenic Systems for Superconducting Magnets, G. Y. RoBINSON, JR., Massachusetts Institute of Technology . 0... F-11 Minimization of Refrigeration Power for Large Cryogenic Systems, M. A. HILAL, Michigan Technological University, and Y. M. EYSSA, University of Wisconsin-Madison . . . . . . . . . . . . . . . . . . Heat Transfer in Helium G-1 Two-Dimensional Heat Transfer to Superfluid Helium, M. A. HILAL, Michigan Technological University .......... G-2 Heat Transfer to Helium-li in Cylindrical Geometries, S. W. V AN SeiVER and R. L. LEE, University of Wisconsin-Madison G-3 Maximum and Minimum Heat Flux and Temperature Fluctuation in Film-Boiling States in Superfluid Helium, H. KoBAYASHI and K. YASUKÖCHI, Nihon University ................. 0.............. G--4 Transient Heat Transfer in Boiling Helium-! and Subcooled Helium-li, P. SEYFERT, G. CLAUDET, and M. J. McCALL, Centre d'Etudes Nucleaires de Grenoble, and R. AYMAR, Centre d'Etudes Nucleaires de Fontenay aux Roses . . . . . . . . . . . . . . . . . . . . . . . . . . . . G-5 Heat Transfer Measurement With a Small Superconducting Coil Subjected to Transient and Quasistatic Heating at Temperatures between 1.8 and 4.2 K, D. GENTILE, Centre d'Etudes Nucleaires de Saclay, and W. V. HASSENZAHL, Los Alamos Scientific Labaratory .............................................. G-6 Heat Transfer of Helium in a Pipe With Suction, L. L. VAsiLIEV, G. I. BosROVA, and L. A. STASEVICH, The Luikov Heat and Mass Transfer Institute . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . G-7 Heat Transfer and Helium Replenishment in Cabled Conductor Cooling Channels, P. F. MICHAELSON, R. QuA Y, and R. F. KOENIG, General Electric Company, and P. L. WALSTROM and J. S. GoooARD, Oak Ridge National Labaratory . . . . . . . . . . . . . . . . . G-8 Measurements of Heat Transfer and Helium Replenishment in Long Narrow Channels, R. E. ScHWALL, F. J. RELES, and J. P. HEINRICH, Intermagnetics General Corporation . . . . . . . . . . . . . . . . . G-9 Vapor Locking and Heat Transfer under Transient and SteadyState Conditions, C.-J. CHEN, S.-T. WANG, and J. W. DAWSON, Argonne National Labaratory................................. G-10 Forced Two-Phase Helium Cooling of Large Superconducting Magnets, M. A. GREEN, W. A. BuRNS, and J. D. TAYLOR, Lawrence Berkeley Labaratory . 0 ••••• 0 0 0. 0 0 • 0 0 ••••••••••••••••••••••••• • • • • • • • 300 308 317 326 335 342 350 358 363 372 378 385 393 398 406 412 420 Contents Heat Transfer H-1 H-2 H-3 H-4 H-5 H-6 H-7 H-8 Contact Heat Transfer in Solid Cryogens, B. I. VERKIN, R. S. MIKHALCHENKO, V. F. GETMANETS, and L. G. GoNCHARENKO, Academy of Seiences of the Ukrainian SSR . . . . . . . . . . . . . . . . . . . . Digital Computer Simulation of Voidage in a Regenerator, J. B. HARNESSand P. E. L. NEUMANN, University of Bradford........ Simulation of Cooldown Underneath Large Cryogenic Storage Tanks, M. H. SEELAND and K. D. TIMMERHAUS, University of Colorado . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Transient Pool Boiling of Liquid Helium Using a TemperatureControlled Heater Surface, P. J. GIARRATANO and N. V. FREDERICK, NBS Thermophysical Properlies Division . . . . . . . . . . . Heat Transfer during Subcooled Hydrogen Boiling, B. I. VERKIN, Yu. A. KIRICHENKO, and N. M. LEVCHENKO, Academy of Seiences of the Ukrainian SSR................................ Observation of Bubble Formation Mechanism of Liquid Nitrogen Subjected to Transient Heating, 0. TsuKAMOTO and T. UYEMURA, Yokohama National University, and T. UYEMURA, Tokyo University . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Natural Convection Heat Leak in Supercritical Hydrogen Tanks, A. J. BARRETT, Beech Aircraft Corporation . . . . . . . . . . . . . . . . . . . . Techniques for Reducing Radiation Heat Transfer between 77 and 4.2 K, E. M. W. LEUNG, R. W. FAST, H. L. HART, and J. R. HEIM, Fermi National Accelerator Laboratory . . . . . . . . . . . . . . . . . . 431 438 446 455 467 476 483 489 Flow Phenomena J-1 J-2 J-3 J-4 J-5 K-1 K-2 Experience with an Orifice Flow Meter lnstalled in a Helium Refrigerator, J. W. DEAN and W. F. STEWART, Los Alamos Scientific Laboratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Some Observations of a Free Jet Phenomenon in a 90° SharpEdge lnlet Geometry, R. C. HENDRICKS, NASA, Lewis Research Center..................................................... Experimental Study of Flow Instabilities in Forced Helium Cooling Channels, Y. MATSUBARA, A. SuGAWARA, and K. YASUKÖCHI, Nihon University................................ Acoustic Oscillation Phenomena in Low-Velocity Steady Flow with Heating, J. A. LIBURDY, Clemson University, and J. L. WoFFORD, Arkansas Power and Light Co. . . . . . . . . . . . . . . . . . . . . . Control of Pressurized Superfluid Helium-li: Application to Loss Analysis, M. X. FRAN(:OIS, Laboratoire d'Aerodynamique, and J. C. LoTTIN and J. PLANCOULAINE, Centre d'Etudes Nucleaires de Saclay..................................................... Liquefaction and Refrigeration Thermodynamic Optimization of Helium Liquefaction Cycles, R. H. HUBBELL, Arthur D. Little, Inc., and W. M. ToscANO, FosterMiller Associates............................................ A 50-Liters/hr Helium Liquefier for a Superconducting Magnetic Energy Storage System, P. J. KERNEY and D. A. McWILLIAMS, CTI-Cryogenics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 500 506 521 528 541 551 563 X K-3 K--4 K-5 Contents A Closed-Cycle Thermally Activated Regenerative Cryogenerator, J. B. HARNESSand P. E. L. NEUMANN, University of Bradford . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A Study of Refrigeration for Liquid-Nitrogen-Cooled Power Transmission Cables, R. C. LoNGSWORTH, Air Products and Chemicals, Inc., and K. F. ScHOCH, General Electric Company . . . The Effect of Temperature on lmpurity Adsorption from Hydrogen on Activated Carbon and Silica Gel, W. C. KRATZ, Air Products and Chemicals, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 574 585 597 Solid and Fluid Properties L-1 Effect of Nitrogen Presence on Solid Solubility of Normal Paraffins in Liquid Ethane, J. D. HorroVY, W. L. CHEN, J. P. KoHN, and K. D. LUKS, University of Notre Dame.............. L-2 Solubility of Solid tert-Butyl Mercaptan in Liquid Methane and an LNG Mixture, G. P. KuEBLER and C. McKINLEY, Air Products and Chemicals, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L-3 Prediction of C02 Solubility in Light Hydrocarbon Mixtures at Low Temperatures, R. J. J. CHEN, V. W. LIAW, and D. G. ELLIOT, DM International, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L--4 Melting Point Data for Freeze Protection in Natural Gas Liquids (NGL) Plants Using Methanol Dehydration, R. L. HoRTON and G. C. DYSINGER, Phillips Petroleum Company.................. L-5 Measurement of Isochoric P-V-T Behavior of a Nominal95mol.%-Methane-5-mol.%-Propane Mixture from Near-Ambient to Cryogenic Temperatures, K. ARAI and R. KoBAYASHI, Rice University . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L-6 Phase Behavior of Three Hydrogen-Containing Ternary Systems, M. YoRIZANE, S. YoSHIMURA, and H. MAsuoKA, University of Hiroshima, I. FUNADA: Kobe Steel Ltd., and C.-T. Fu and B. C.-Y. Lu, University of Ottawa . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L-7 Vapor Pressure and Heats of Vaporization and Sublimation of Liquids and Solids of Ioterest in Cryogenics Below 1-atm Pressure, G. N. BROWN, JR. and W. T. ZIEGLER, Georgia Institute of Technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . L-8 Simple and Generalized Equation of State for Vapor-Liquid Equilibrium Calculations, T. ISHIKAWA, W. K. CHUNG, and B. C.-Y. Lu, University of Ottawa............................... L-9 Reduced Volumetrie Expansion and Isothermal Compressibility Factor Plots, R. C. HENDRICKS, NASA, Lewis Research Center . . . L-10 The Construction and Use of Exergy Diagrams, W. 0. DALY and J. B. IIARNEss, University of Bradford . . . . . . . . . . . . . . . . . . . . . . . . L-11 Densities and Dielectric Constants of LPG Components and Mixtures at Cryogenic Storage Conditions, R. T. THoMPSON, JR. and R. C. MILLER, University of Wyoming . . . . . . . . . . . . . . . . . . . . . L-12 Dielectric Properlies of Liquid Hydrogen, V. V. PASHKOV and M. P. LoBKO, Ukrainian Academy of Seiences . . . . . . . . . . . . . . . . . 609 616 620 629 640 654 662 671 682 693 698 709 Contents xi Cryogenic Applications-LNG M-1 M-2 M-3 M-4 M-5 World Market for LNG Trade-Present and Future, B. M. GIBSON, Air Products and Chemicals, Inc. . . . . . . . . . . . . . . . . . . . . . LNG Project Development: Shipping and Terminal Considerations, V. V. STAFFA and D. K. JHAVERI, Tennessee Gas Transmission Company.................................. Marine Transportation of LNG at Intermediate Temperature, C. P. ßENNETT, University of Manitoba . . . . . . . . . . . . . . . . . . . . . . . . . . Explosion Hazards of LNG and LPG Carriers during Transport, M. M. KAMEL and A. KHALIL, Cairo University................ Purification and Cryogenic Separation of SNG Produced from Coal, A. A. CASSANO, T. C. LI, J. C. TAo, and T. R. TsAO, Air Products and Chemica/s, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 715 730 751 757 763 Cryogenic Applications-Space Technology N-1 N-2 N-3 N-4 A Portable 3 He Cryostat for Studies in Astrophysics, A. SHERMAN and 0. FIGUEROA, Goddard Space Flight Center...... Mechanism of an Active Phase Separator for Space Applications, H. D. DENNER, G. KLIPPING, I. KLIPPING, K. LÜDERS, J. MENZEL, and U. RuPPERT, Freie Universität Berlin . . . . . . . . . . . . . Progress on the Development of a 3- to 5-Year Lifetime Stirling Cycle Refrigerator for Space, A. SHERMAN and M. GASSER, Goddard Space Flight Center, M. GoLDOWSKY, North American Phi/Ups Corporation, G. ßENSON, Energy Research and Generation, Inc., and J. McCoRMICK, Mechanica/ Techno/ogy, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Superfluid Helium Experiment for Spacelab 2, P. MASON, D. COLLINS, P. COWGILL, D. ELLEMAN, D. PETRAC, M. SAFFREN, and T. WANG, Jet Propulsion Labaratory....................... 775 783 791 801 Cryogenic Applications-Resource Utilization 0-1 0-2 0-3 Helium: lts Past, Present, and Future, E. F. HAMMEL, Los Alamos Scientific Labaratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Liquid Hydrogenas an Automotive Fuel, W. F. STEWART, Los Alamos Scientific Labaratory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Liquid Nitrogenas an Energy Source for an Automotive Vehicle, M. V. SussMAN, Tufts University . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 810 822 831 Indexes Author Index Subject Index 839 843 FOREWORD The Cryogenic Engineering Conference celebrated its Silver Anniversary at the 1979 Conference held at Madison, Wisconsin. For many it provided an opportunity to reminisce about the first Cryogenic Engineering Conference convened at the National Bureau of Standards in Boulder, Colorado and also about the many following conferences and advances that had been reported at these conferences. It is difficult to realize that the first Cryogenic Engineering Conference was held before the advent of multilayer insulation, the space age, large-scale LNG Operations and superconductivity applications. The evolution of these activities has been carefully recorded in past volumes of the Advances in Cryogenic Engineering. · Once again, the Cryogenic Engineering Conference is happy to have had the International Cryogenic Materials Conference cohost this meeting at the University of Wisconsin. Collaboration between these two conferences has proven to be mutually beneficial by providing the cryogenic engineer with an in-depth exposure to materials properties, selection, and utilization to complement the exposure to new applications and design concepts. The papers presented at this joint conference as part of the International Cryogenic Materials Conference will be published as Volume 26 of the Advances in Cryogenic Engineering. Many have contributed to the success of the 1979 Cryogenic Engineering Conference in Madison. The Cryogenic Engineering Conference Board is extremely grateful to R. W. Boom and his staff at the University of Wisconsin for their excellent handling of all the local arrangements and to the University of Wisconsin for serving as hosts for this Silver Anniversary meeting of cryogenic specialists from all over the world. The assistance of the many dedicated workers in the cryogenic field who have once again contributed to the reviewing of the final manuscripts for this volume is also gratefully acknowledged by the Cryogenic Engineering Conference Board and the editor. The list of all those individuals who have assisted in the many important tasks involved in completing the manuscripts for publication continues to grow Ionger with each volume in this series, and any attempt to acknowledge individual contributions in this limited space would not do justice to them. However, Special recognition, as in past years, must be given to Mrs. Elva R. Dillman from the University of Colorado for her continued assistance to the editor and her attention to all the details involVed in the preparation of the final manuscripts for this and past volumes. Her devotion to this activity has been exemplary. This series has traditionally recognized individuals who in some way have contributed significantly in extending the frontiers of cryÖgenic engineering or have provided noteworthy assistance to the Cryogenic Engineering Conference. In the spirit of this tradition this series recognizes another individual who has been an active and eftective leader in the application of cryogenic engineering to a wide variety of engineering problems and has ably supported the Cryogenic Engineering Conference. Accordingly, Volume 25 of the Advances in Cryogenic Engineering is dedicated to H. 0. McMahon, retired director of Artbur D. Little and recipient of the fourth Samuel C. Collins Award presented on August 23, 1979. :di RUSSELL B. SCOTT MEMORIAL AWARD The objectives of this award are to provide increased recognition for the recipients in the scientific community and to provide an incentive for high er quality in both oral and written presentations at future Cryogenic Engineering Conferences. The winners of the Russell B. Scott Memorial Award for the outstanding papers presented at the 1977 Cryogenic Engineering Conference, as announced by the Awards Committee, are as follows: In the cryogenic engineering research category, T. R. Dali and J. C. Chato of the University of lllinois are recognized for their paper "Effects of Natural Convection on Heat Transfer in Porous Cryogenic Insulations." In the application of cryogenic engineering category, C. J. Mole, P. W. Eckels, H. E. Haller, III, M. A. Janocko, S. A. Karpathy, D. C. Litz, E. Mullan, P. Reichner, and Z. N. San j ana of the Westinghouse Research and Development Center, and D. W. Deis of the Lawrence Livermore Laboratory, and M. S. Walker of Intermagnetics General Corporation are acknowledged for their paper "A Superconducting 0.54-MJ Pulsed Energy Storage Coil." The Cryogenic Engineering Conference extends its congratulations to all of these award-winning authors. xiii SAMUEL C. COLLINS AWARD The Samuel C. Collins Award for outstanding contributions to cryogenic technology was presented to Howard 0. McMahon at the 1979 Cryogenic Engineering Conference held at the University of Wisconsin. McMahon is a retired president of Artbur D. Little, Inc., and a member (and form er chairman) of the board of Cryogenic Technology, Inc., now Helix Technology Corporation. The Collins Award was established in 1965 by the Cryogenic Engineering Conference to honor Samuel C. Collins, professor of mechanical engineering at Massachusetts Institute of Technology, who in 1946 invented the first practical helium liquefier. Dr. Collins, retired from MIT in 1964, received the Award in 1965 and is now associated with the U. S. Naval Research Laboratory, where he is active in the development of a helium Iiquefier for shipboard use. The Collins Award was presented to Professor Klaus D. Timmerhaus, University of Colorado, Boulder, in 1967 and to Dr. Edward F. Hammel, Los Alamos Scientific Laboratory in 1973. Howard 0. McMahon, a native of Killman, Alberta, Canada, received bis B.A. and M.A. degrees from the University of British Columbia and bis Ph.D. from MIT. He joined Artbur D. Little, Inc. in 1943 and worked with Dr. Collins on the development of the first helium liquefier. His contributions to cryogenic technology included the McMahon rectifying column packing, air separation heat exchangers, and the coinvention of the Gifford-McMahon cryogenic refrigerator. The latter made possible the development of efficient, highly reliable closed-cycle cryogenic refrigerators for temperatures below 20 K, which are now used in a variety of industrial and research applica~ions. He has received the Edward Longstreth Medal of the Franklin Institute and the Frank Forrest Award of the American Ceramic Society. Dr. McMahon is the author of many technical papers and holds 22 patents on a wide variety of inventions, many in the field of cryogenic engineering. Howard 0. McMahon, recipient of the Samuel C. Collins Award, August 21, 1979. xiv 1979 CRYOGENIC ENGINEERING CONFERENCE BOARD R. W. Fast, Chairman . . . . . . . . . . . . . . R. F. Barron . . . . . . . . . . . . . . . . . . . . . . R. W. Boom . . . . . . . . . . . . . . . . . . . . . . M. B. Clapp . . . . . . . . . . . . . . . . . . . . . . . D. B. Crawford, Vice-Chairman . . . . . T. M. Flynn, Program Chairman . . . . . C. D. Henning . . . . . . . . . . . . . . . . . . . . . M. J. Hiza, Jr. . . . . . . . . . . . . . . . . . . . . . J. E. Jensen· . . . . . . . . . . . . . . . . . . . . . . . E. R. Lady . . . . . . . . . . . . . . . . . . . . . . . . R. C. Longsworth . . . . . . . . . . . . . . . . . . Fermi National Accelerator Labaratory Louisiana Tech University University of Wisconsin, Madison Chicago Bridge and Iron Company Puliman Keliogg National Bureau of Standards University of California, Lawrence Livermore Labaratory National Bureau of Standards Brookhaven National Labaratory University of Michigan Air Products and Chemicals, Inc. ex officio B. W. Birmingham . . . . . . . . . . . . . . . . . R. P. Reed . . . . . . . . . . . . . . . . . . . . . . . . K. D. Timmerhaus, Editor, Advances in Cryogenic Engineering . D. A. Belsher, Administrator . . . . . . . . National Bureau of Standards National Bureau of Standards University of Colorado National Bureau of Standards A WARDS COMMI'ITEES RUSSELL B. SCOTI MEMORIAL A WARD COMMITIEE M. J. Hiza, Jr., Chairman . . . . . . . . . . . R. F. Barron . . . . . . . . . . . . . . . . . . . . . . E. R. Lady . . . . . . . . . . . . . . . . . . . . . . . . K. D. Timmerhaus . . . . . . . . . . . . . . . . . National Bureau of Standards Louisiana Tech University University of Michigan University of Colorado S. C. COLLINS AWARD COMMITIEE R. W. Fast, Chairman . . . . . . . . . . . . . . D. B. Crawford . . . . . . . . . . . . . . . . . . . . T. M. Flynn . . . . . . . . . . . . . . . . . . . . . . . K. D. Timmerhaus . . . . . . . . . . . . . . . . . Fermi National Accelerator Labaratory Puliman Kellogg National Bureau of Standards University of Colorado ACKNOWLEDGMENTS The Cryogenic Engineering Conference Board is deeply grateful for the support which the following organizations have given to the 1979 Cryogenic Engineering Conference. The Aerospace Corporation AIRCO, Inc. Air Products and Chemicals, Inc. Ball Corporation Chicago Bridge and Iron Company General Electric Company Helix Technical Corporation, CTI-Cryogenics Helix Process Systems Division National Bureau of Standards, Boulder Laboratories Puliman Kellogg Silbrico Corporation Union Carbide Corporation, Linde Division University of Colorado University of Wisconsin, Madison A-1 SUPERCONDUCfiNG MUD MAGNET ENGINEERING PROGRAM P. G. Marston, A. M. Dawson, D. 8. Montgomery, and J. E. C. Williams Francis Bitter National Magnet Laboratory Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCfiON MHD has the potential to become the first large-scale commercial application of superconductivity. Recent accomplishments of the DOE/MHD development program have shown this fossil-fueled power generation technique to be capable of higher efliciencies and lower-per-kilowatt-hour cost than any other near-term technology. A significant scale demonstration of all critical items is scheduled to begin in late 1979 at the DOE Component Development and Integration Facility (CDIF) presently under construction in Butte, Montana. A sequence of progressively more complex ftow train configurations will be demonstrated through 1986. Commercial scale system and optimization studies have already identified component requirements. The next step will be operation, during the next decade, of a system large enough to be conservatively scaled to commercial size. These systems will require large, complex, and costly superconducting magnets operating at fields of at least 6 T. The MIT Francis Bitter National Magnet Labaratory has been designated as the DOE magnet program field oflice to assist with the creation and management of a national program of superconducting MHD magnet technology development. This activity includes conceptual design and subsequent contract management for a nurober of large magnets destined for use in a variety of MHD experimental facilities. Figure 1 shows the relative size and weight of superconducting magnets presently under consideration. The smallest of these is the U-25 Bypassmagnet now in use at the Soviet High Temperature Institute MHD pilot facility near Moscow. The next larger unit represents the superconducting magnet under construction for CDIF. The two largest magnets represent sizes identified for small commercial and base-load power generation plants. The intermediate-sized unit is that presently proposed for the Engineering Test Facility (ETF). A five-year program of technological and industrial development willlead to the selection of a design for the ETF magnet by late 1983 and construction by the late 1980's. This willlead ultimately to the construction of a commercial scale magnet for a coal-burning MHD/steam power plant operating in a utility environment. 1 P. G. Manton, A. M. Dawson, D. B. M011tcomery, IUid J. E. C. Williams MAGNET APPLICATION MAGNET FIELD STRENGTH (T) ELEVATION VIEW STORED MAGNETIC ENERGV ( ALL VIEWS DRAWN TO SAME SCALE l MJ WARM BORE SIZE CRYOSTAT SIZE MAX. 00/LENGTH MAGNET WEIGHT INLET/EXIT ( metersl (lonnes) ( metersl BASELOAD 1000 MWe MHD 6 16000 ------------ 2 9/6 5 16/37 -7500 12 5/24 3000 SQ * 250 MWe MHD ETF (AVCOl 6 6000 2 1/3 9 SQ 6 1500 I 5/2 2 8 10 5/14 9 0.710 9 4 3/63 150 2 3/44 28 SO. 1500 CDIF iGEl U25 BYPASS (ANLl 6 IBO SQ 5 32 0 5 10 0.4/0 67 DIA '---.J..._J SCALE-METERS Fig. 1. MHD magnet comparison. The technology development program is dynamically integrated with the magnet construction management in order to achieve the following: identify design and manufacturing techniques for commercial scale units; identify failurc modes, safety and risk considerations; define evaluation and success criteria; predict costs; provide the fundamental engineering data base and design tools; and perform verification testing and modeling. Overall program network analyses have been developed. These show parallel paths pursued in high-risk areas and component development on a progressively increasing scale. Analytical tools are being refined and documented. A high degree of expertise has been achieved via the use of industrial, university, and national laboratory subcontractors, and by periodic reviews, conferences, seminars, workshops, and summer teaching courses. MHD MAGNETS Of the many configurations capable of providing suitable magnetic field distributions, the two basic concepts of most interest are the circular and the reetangular saddle configurations. Figure 2 provides an artist's conception of a base..Joad size MHD magnet, and Fig. 3 shows the CDIF size magnet system. Figure 4 presents comparative cross sections of the two designs. The major components ar4 given in the following sections. Willding Substracture. Each superconducting wire in the winding rn\ISt be in intimate contact with liquid helium for cooling, must be weH insulated to withstand transient voltages during emergency discharge, and must be weH supported against Superconducting MHD Magnet Engineering Program 3 Fig. 2. Base-load magnet using circular saddle design. the !arge electromagnetic Lorentz forces. Most of the Iarge-scale magnet designs identified today accomplish at least the mechanical support function via some form of mechanical substructure which subdivides the winding into a number of mechanical regions and thus Iimits cumulative Ioads on conductors and insulation. Superstructure. This is the outer force-containment structure which supports the cumulative electromagnetic load. The total bursting force (y direction, Fig. 4) for a base-load size magnet is approximately 5000 ton/m of axiallength. Fig. 3. CDIF superconducting magnet using reetangular saddle design. 4 P. G. Marston, A. M. Dawson, D. B. Montgomery, and J. E. C. Williams z types ol section can be used) WINDING AREA V ELEMENT IN BENDING TENSION MEMBERS Fig. 4. Comparative cross sections of reetangular and circular saddle designs. HeUum Vessel. The containment vessel for the liquid helium coolant may be external to the superstructure and therefore self-supporting. But in the larger units, cost considerations willlikely require the use of the superstructure to support the internal pressure Ioads of the helium vessel. This and the above elements constitute the "cold mass." Cold·Mass Supports. A very strong support system must minimize thermal conduction to the cold rnass and also permit thermal contraction of the cold mass relative to the room temperature support structure. A technique suitable for the smaller magnets can be seen Fig. 3 and consists of eight glass-epoxy links which are angled and capable of rotating in such a way that thermal motion is accommodated without relaxing link tension. Support schemes for larger magnets are being studied and willlikely incorporate pivoting support posts. 5 Superconducting MHD Magnet Engineering Program Vacuum Vessel. Radiation heat loads to the helium temperature cold mass are limited by vacuum insulation with an intermediate radiation shield maintained near liquid nitrogen temperature. Multilayer insulation is incorporated on either side of the radiation shield. Ancillary Systems. Operation of the magnet requires a large helium refrigerator/liquefier to cool down, fill, and automatically maintain the liquid helium level in the cryostat for long periods of time. A block diagram of the cryogenic system is shown in Fig. 5. The energizing, instrument, and protection systems must allow a 1-hr discharge-charge cycle for fast channel replacement and must also provide a means for removing most of the stored energy in the magnetic field in the event of system failure. This is normally accomplished by use of a dump resistor permanently connected across the magnet current leads. When the energizing circuit is opened, the magnet current decays according to the Lmagnet! Rdumpres. time constant; nearly all the stored energy is dissipated into the resistor. UTILITY VACUUM PUMP N2 VENT COOLDOWN EXCHANGE He GAS STORAGE COMPRESSOR [ ~r-- _-;>_ _-;>. F~LL LN2 STORAGE N2 VENT WCONNECTOR - VENT He REFRIGERATOR COLDBOX ..9- _-;>_ ~~ t ~ n LEADS FILL cON~ ~ N2L LHe STORAGE MAGNET -·· Fig. 5. Cryogenic system block diagram. 6 P. G. Marstoa, A. M. Dawma, D. B. Moatxo~~~ery, ud J. E. C. WDams TECHNOLOGY DEVELOPMENT Superconductors and structures have received the greatest amount of attention to date, in part because these inftuence most strongly the design of immediate (CDIF, CFFF, Stanford) and intermediate (ETF, CDP) magnet systems. Supercondudor Study. Superconductor studies have been conducted both by this laboratory and industry. This laboratory has developed the minimum propagating zone (MPZ) theory. With this theory it is possible to calculate the length of a conductor which will "quench" from a given thermal energy perturbation and determine whether or not that quench normalzonewill recover to a superconducting state or whether the quench will propagate throughout the entire magnet [ 1]. This work has been extended by lwasa and Apgar to a study of the transient heat transfer to the liquid helium layer under film boiling conditions. An equation was derived in that study for the transient heat transfer rate given by eJ d(} q,((}) = q.((}) + a((}) dt (1) where a(8) is the effective heat capacity of the vapor layer. The term containing a(8) is strongly dependent on surface conditions. Because of this term, transient heat transfer in the film boiling region is much higher in the heating of a conductor, and lower in the cooling than steady-state heat transfer; thus, heat transfer is hysteretic with temperature cycling. This function determines quantitatively the vapor-layer thickness as a function of temperature and thereby provides a guideline for sizing cooling channels irr a superconducting winding. Sindair and Iwasa [3 ] have also studied the time response of a superconductor to a thermal perturbation. The duration of a perturbation was shown both analytically and experimentally to have a large inftuence on quench energy. Brief, high-energy perturbation can be dissipated without quench while a significantly lower energy perturbation over a Ionger time period might Iead to quench. This work is now being extended to consider the spatially dependent case. The MPZ theory outlined above has been applied in the design of both the Stanford and CDIF superconducting magnets, and will be applied in the ETF, CDP, and base-load designs. Frictional heating is a potential source of quench-producing energy input in a superconducting magnet winding. Energy input from the movement between conductor and insulator or conductor and substructure is being studied in this laboratory to identify the most suitable materials pairs for future magnet construction. This work, when complete, will provide another means of predicting a potential perturbation. Study contracts have been awarded to industry to evaluate stability and costs for high-current superconductors of 25-250 kA. Several conductor configurations being considered are shown in Fig. 6. The cost analyses have indicated a slight advantage for a 100-kA conductor but the cost minimum is broad; more detailed studies could push the optimum current as low as 25 kA. Present CDP studies use a 50-kA conductor as a baseline. The stability analyses considered three different effects: eddy current heating due to ftux diffusion, reduction of recovery current due to transverse thermal gradients, and reduction of recovery current due to transverse electrical currents. These are particularly important when the operaring current exceeds the Maddock et al. equal area criterion Iimit [4 ]. Each of these effects reduces the MPZ energy. 7 Superconducting MHD Magnet Engineering Program ,-/ y ~- ' /- ~ Cable superconduc tor I Fill er 1 rmm ~ ~~ Wl15 .42 I '-- Superconducting cable Copper submale Supercond uctor strand Copper ~ -=-=r strand~ ~ Co~per 5.42 mm substratc Separated substrate-Multiple cables, single channel 1.45mm lntegrated substrate conductor 12.5 kA modules ~1mm ~ Coppersubstrate Superconducting cable sol dered 10 Substra te Fig. 6. Suggested high-current conductor configurations. Single integrated substrate conductor Preliminary information has been developed regarding the effect of superconducting-to-normal transition time, stabilizer and solder properties which give rise to transverse thermal gradients, and the effect of stabilizer thickness. Experimental verification of these predictions and related analyses is scheduled during fiscal year 1980. ICCS. Recent studies of the behavior of internally cooled cabled superconductors (ICCS) have demonstrated two important characteristics. The first is the demonstration that frictional heat generated by Lorentz force compaction of the cable is small and does little to degrade performance. The second is that even small thermal perturbations generate high turbulence and local heat transfer so that very little pumping or forced cooling is required.* Parameters affecting performance are being studied analytically by Arp at the National Bureau of Standardsand experimentally by Hoenig, Iwasa, and Becker at FBNML. * Hence the change in terminology from "force cooled" to "internally cooled." 8 P. G. Manton, A. M. Dawson, D. 8. Montgomery, and J. E. C. Williams Struetures. The electromagnetic force COntainment structure has two elements: a substructure which is interposed with the conductors in such a way that the winding is divided into a number of mechanical regions within which cumulative Ioads of conductor and insulation can be limited to safe values; and a superstructure which surrounds the winding/s,bstructure composite and contains overall magnetic forces. Many conductor--substructure--superstructure combinations are possible and their impact on final iQstalled cost creates a complex matrix of options to be evaluated. Of particular importance is the impact of shipping size and weight limitations on the need for on-site construction and test. The CDIF, CFFF, and Stanford magnets represent three quite different structural design concepts. Relative costs, fabricability, reliability in a utility environment, and scalability will be compared during their design, construction, and test. MAGNET PROCUREMENT CDIF Superc:ondu~g Magnet. The CDIF magnet, shown in Fig. 3, uses a glass-reinforced epoxy (NEMA G-10) substructure consisting of grooved subplates wherein each conductor is supported in its own groove. This approach reduces cumulative Ioads and frictional heating to an absolute minimum. It virtually eliminates any need for mechJlnical strengthin the conductor and might permit the use of a flexible cable which could be wound on-site very easily. The design eliminates the need for conductor insuhttion and provides a winding subunit which is handled and assembled easily and has good mechanical integrity, strength, and precision. Preliminary tests suggest tbat the helium coolant channels may be eliminated. This simplification will be verified in the test facility magnet now under construction. An extensive study of the low-temperature mechanical properties of commercially available G-1 0 has shown the material to be satisfactory but requires careful quality assurance monitaring of the manufacturer's process. Manufacturing studies indicate that a comparable composite could be molded, giving superior mechanical properties for the intended use and costing about one-quarter that of machined G-10. The CDIF superstructure is a straighttorward system of crossbeams and tension members. The componeots are of 304LN type stainless steel weldments. Submerged arc welding appears to be adequate but this has yet to be verified. Sealeup. A scaling and cost study to ETF and CDP size is included in the CDIF work statement [5]. A comparable study by A VCO-Everett Research Labaratory considers a magnet havbig similar winding configuration but quite different substructure and superstructure of stainless steel and aluminum, respectively. The AERL concept incorpotrates individual helium vessels for each winding half, thus enabling the first order of modularization to deal with shipping size Iimits. Several other reetangular saddle eonfigurations have been proposed which will be the subject of study co~tracts during fiscal year 1980. StaDlord Mapet. The Stanford magnet design concept uses semicylindrical aluminum subplates in which the number of conductors per groove is limited to that giving safe cumulative compressive Ioad on the turn-to-turn insulation. The superstructure consists of circular aluminum ring girders. The highly conductive and closely coupled substructure will modify quench behavior to the degree that such a design could be inherently fail-safe without the need for a dynamic energy-dumping system. Superconducting MHD Magnet Ellgineering Program 9 Certain problems of scaling are being addressed by General Dynamics/Convair Division who have proposed that the cylindrical substructure be fabricated from a series of longitudinal staves. This concept has thus been dubbed "cask." As in the case of the rectangular-saddle configuration, additional circular-saddle studies are anticipated. CFFF Coai-Fired Flow Facility Magnet. This magnet is being designed and built by Argonne National Laboratory. lt incorporates important similarities but also important differences from the Argonne-built U-25 Bypass magnet presently operating at the Soviet High Temperature Institute. Both magnets have individually banded circular-saddle winding layers with micarta filler material in the winding bore. In the U-25B, the banding is of very tightly wound stainless steel. The radial forces are supported by a strong, thick stainless steel bore tube. Loads are transmitted through the micarta fillers and bands. In the CFFF magnet, the banding is of aluminum and is used solely as a manufacturing aid to secure individuallayers as the winding is "built up." The radial forces are supported by external curved beams cast from stainless steel and connected through aluminum tension links. Support of the axial force components is also quite different. In the U-25B, a very large portion of this Ioad is communicated to flanges on each end of the bore tube via transfer structure tightly fitted at assembly between the saddles and flanges. In the CFFF magnet, these Ioads are taken directly by the conductors. It is anticipated that Argonne will perform scaleup studies for the CFFF concept. Characteristics of the three magnets presently under construction are listed in Table I. Table I. Characteristics of the CFFF, CDIF, and Stanford Snpercondncting Magnets Design characteristics CFFF CDIF/SM Stanford SM SC SC SC Peak on-axis field,* T Inlet bore dimensions, m Outlet bore dimensions, m Active field length, m Overall dimensions, m Overalllength, m Total weight, tons Stored energy, J Dipole moment, A-m 2 6 0.8diam. 1.0 3 4.9 X 4.1 6.4 172 168 X 106 6 0.98 X 0.78 0.98sq. 3 4.3 X 4.3 6.9 180 184 X 106 1 X 108 7.3 0.55 diam. 0.55 diam. 1.5 3.8 X 3.8 4.5 70 79 X 106 1.2 X 108 Conductor JA, A/m2 Heat tlux, W /cm 2 Operating current, A Operating power, MW 2.0 X 107 0.135 4,000 1.65 X 107 0.4 6,000 2.08 X 107 0.7 5,000 ANLt 8 GE:t: 5.4 GD§ Type Estimated cost, $106 Contractor * Field horizontal in all cases. t Argonne National Laboratory. :j: General Electric Company Energy Systems Programs Division. § General Dynamics Corporation, Convair Division. 10 P. G. Ma-ston, A. M. Dlnnon, D. B. Moatpmery,lllld J. E. C. W&.ms ANCILLARY STUDIES Cryogenic Support Systems. A preliminary design and cost study by Cryogenic Consultants, Inc. has identified helium refrigerator/liquefier systems in the range of 100-200 liters/hr of liquid-generating capacity and 400~800 W of refrigeration as the range of interest for 50-kA commercial scale systems. Power and Protection. A preliminary design and cost study for energizing, energy dumping, control, and protection systems is in process by A. Kusko, Inc. Shipping Study. Belding Corporation, the firm that transported the U-25B magnet from Argonne to Moscow, has completed preliminary study of the overall problern of shipping size and weight limitations. This information will help establish the degree of modularity required and the relative amount of factory vs. on-site construction required as a function of MHD generator size. Isotensoidal Force Containment Concept. An advanced concept intended to give a minimum cost baseline is being studied by Batteile Columbus Laboratories. In this design, an internally cooled cabled conductor is used in such a way that its stainless steel jacket is a major structural element. The saddle configuration is isotensoidal and the axial Ioads are taken by the individual conductors (jackets). The transverse forces on the straight sections of the windings are contained by banding, which also has an isotensoidal configuration. Cost Studies. Preliminary cost estimates have been made for several magnets. Credible scaling factors have been determined, but considerable additional effort and design refinement are required to achieve a reasonable degree of cost credibility. Properfies of Materials. When one considers the long-term operation in a utility environment of thousands of tons of exotic alloys operating highly stressed at near absolute zero temperature, the inadequacy of existing data on properties of materials becomes apparent. The degree of conservatism which must therefore be imposed on allowable stress Ievels is also very costly. Several contracts with the National Bureau of Standards (Boulder, Colorado) have been initiated to investigate both structural and conductor materials and composites. Facility. The existing facilities at FBNML are being upgraded to permit a variety of significant-scale structural tests intended to simulate the actual operating conditions of proposed designs and to determine their eiementaland composite behavior. One such test scheduled for fiscal year 1980 will examine the structural behavior of internally cooled cabled superconductors. The experimental configuration shown in Fig. 7 will permit a magnetically induced tension in the conductor jacket of up to 100,000 psi in a configuration simulating actual operating conditions. The dimensions and support are configured suchthat the ratio of the stress intensity (tension) in the large radius region to the stress intensity (compressive) in the small radius region is approximately unity. This test will simulate conditions for the isotensoidal design and will also give valuable information for high-field fusion Tokamak coils using niobium-tin (Nb3 Sn). By controlling the cable pitch and thus the relative strain of the conductor and jacket, the stainless steel jacket can be operated at its stress limit rather than the strain Iimit of the Nb 3 Sn conductor. This could reduce the amount of structural material required in such systems by a significant factor. Superconducting MHD Magnet Engineering Program 11 ·~ \ I' ~) "' \ I Fig. 7. Test facility magnet. Development of lower-cost and superior GRP composites is also anticipated to begin in fiscal year 1980. Scale modeling and verification testing of critical system elements is scheduled to begin in mid-fiscal-year 1980. Computer Codes. Existing computer codes for field, force, and stress are being refined and put into easily usable and self-consistent format. These will be fully documented and ultimately available for general use. Subroutines for cost analysis will be developed. CONCLUSIONS The synergism of the parallel procurement and technology development activities permits a dynamic interaction which is a vital part of the national program. The success of the program to date is best illustrated by the fact that only a few years ago large superconducting magnets were considered the pacing technology for MHD power generation. Today there is virtually total confidence in their ultimate success and also in the fact that by the mid-1980's the technology and industrial basewill be adequately in place to move smoothly into what seems certain to be an important new international energy market. REFERENCES 1. 2. 3. 4. 5. M. N. Wilson and Y. lwasa, Cryogenics 18(1):17 (1978). Y. lwasa and B. A. Apgar, Cryogenics 18(5):267 (1978). M. L. Sindair and Y. lwasa, IEEE Trans. Magn. Mag-15(1):347 (1979). B. J. Maddock, G. B. James, and W. T. Norris, Cryogenics 9:261 (1969). J. L. Zar, "A Modular Design for a Superconducting Magnet for the ETF and larger MHD generators," presented at 18th Symposium Engineering Aspects of Magnetohydrodynamics, Butte, Montana, June 1979. A-2 IMPACf OF IDGH-CURRENT OPERATION ON THE COST OF SUPERCONDUCfiNG MAGNET SYSTEMS FOR LARGE-SCALE MHD APPLICATIONS* R. J. Thome,t R. D. PUisbury,t H. R. Segal, and B. 0. Pederson Magnetic Corporation of America Waltham, Massachusetts INTRODUCfiON Reference designs for superconducting magnets for base Ioad magnetohydrodynamic (MHD) plan~s have indicated that they will be substantially larger and more complex in geometry than the largest superconducting magnets now in operation. The cost of many of the components and of some of the steps in fabrication of these magnets may be expected tobe dependent on the current Ievel. A question naturally arises, therefore, as to whether an optimum current Ievel may be found from a cost standpoint. The purpose of this study was to perform a preliminary investigation into this area. The reference design concept C· 2] selected as a baseline is a 26-m-long magnet system with an outside diameter of 9.6 m and an estimated weight of about 1200 tons. Conductor, structure, and dewar represent 14%,69%, and 17% of the total weight, respectively. The magnet consists of an assembly of two reetangular saddle and six racetrack coils nested so as to provide a field variation from 6 to 3.5 T along the 16-m active-field length. Present estimates indicate that a magnet of this size would be suitable for producing roughly 150 MWe from the MHD portion of a combined MHD/conventional plant. The method in this study involved development of a set of cost factors in the general areas of system components, fabrication, and assembly. Components with features expected to be strongly dependent on the current Ievel were studied in sufticient detail to allow their characteristics to be determined for cost-estimating purposes. These included conductor, substructure, the power supply subsystem, and refrigerator/liquefier subsystem. The characteristics (and cost) of components of the superstructure and dewar would not be expected to be a strong function of the current Ievel and were, therefore, assumed to be the same for all cases. Cost estimates for magnet fabrication and system assembly were developed by generating • Study supported by the U. S. Department of Energy, MHD Division through M.I.T. Francis Bitter National Magnet Laboratory's Superconducting MHD Magnet Development Program, DOE Contract No. EX77-A-01-2295. t Present address: Massachusetts Institute of Technology, Francis Bitter National Magnet Laboratory, Cambridge, Massachusetts. 12 Impact of High-Current Operation on the Cost of Superconducting Magnet Systems 13 hypotheticäl manufacturing fiow diagrams, assigning cost elements to the individual steps, and integrating these elements into cost factors dependent on current Ievel (e.g., fabrication of winding and substructure) and cost factors assumed to be independent of current Ievel (e.g., assembly to superstructure, dewar, and support systems). For the purposes of discussion, consider the systemtobe composed of three major subsystems: (1) magnet/dewar, (2) power supply, and (3) cryogenic. MAGNET/OEWAR SUBSYSTEM Conductor The total amp-meterrequirementforthe reference design [2 ]was 1.73 x 109 Am. This may, as a first approximation, be assumed tobe independent of current Ievel. It then determines the total conductor length required once the current Ievel is selected. A baseline configuration for the conductor cross section was chosen and is illustrated in Fig. 1. It consists of a separate, fully transposed cable which is soldered into the substrate during the winding fabrication process. The copper stabilizer is grooved along the wide faces to provide extended heat transfer surfaces. Dimensions for the conductor were based on an operating current at 90% of the short sample current, a simple stability criteria on using an allowable surface heat ftux of 0.6 W I cm2 , and a local surface bearing pressure Iimitation of 10 ksi for the copper. Three conductors were determined for each current Ievel in order to include an allowance for grading of the conductor in a magnet of this scale. This was done by assuming that 30% of the conductor would be sized for 8 T, 45% for 6 T, and 25% for 4.2 T. The weight of the stabilizer was then determined, estimated at a cost of $2/lb and added to the cost of the cabled superconductor with the result that the cost for conductor ranged from $8.2 x 106 at the 10-kA Ievel to $9.4 x 106 at the 250-kA Ievel with the primary variation rising from the stabilizer since the total ampere-meters required was assumed constant. Substructure The internal or substructure concept which was assumed provides a means for cumulative electromagnetic Ioad transmittat around each conductor. Each conductor is inserted in a stainless steel channel which is then closed by a backing strip. The walls and legs of the stainless steel substructure were sized to carry the maximum cumulative Ioad out to the superstructure or to a plane of symmetry where it would be equilibrated. The turn-to-turn insulation consists of a spiral wrap applied to the conductor before insertion in the stainless steel substructural element. Fig. 1. Baseline conductor configuration featuring a separate fully transposed cable of NbTi/Cu composite conductors which is inserted into the copper stabilizer during the winding fabrication process. R. J. Thome, R. D. Pillsbory, H. R. Segal, and 8. 0. Pederson 14 At the lower current levels (i.e., I < 50 kA) the thickness requirements for the channels were such that they could be roll formed. Fabrication of large quantities in this manner was estimated to cost $1.40/lb. At the high current levels (250 kA) the thickness requirements are large fractions of an inch, hence roll forming is probably not feasible. The channel components would most likely be individually rolled then positioned and welded at the corners. The unit price for fabrication in this manner was estimated at $4.05/lb. At intermediate current levels the thicknesses vary between the two extremes, hence a linear variation in unit cost was assumed. Length requirements were based on those generated for the conductor. Total substructural cost estimates ranged from $4 x 105 at the 10-kA level to $4.1 x 106 at the 250-kA level. Superstructure As mentioned earlier, the total amp-meter requirement and, in turn, the overall size of the winding was assumed to be independent of current level. The size of the superstructure, then, also remains constant. The cost for the smaller lighter components was estimated at $4.05/lb as in the case of the heaviest substructure. The drawings of the basic stainless steel structural concept were reviewed with a potential vendor who provided rough estimates of the total cost of the larger components using an equivalent $8.43/lb, which was considered conservative. As a result, the cost of the baseline design superstructure weight of 1.1 x 106 kg was estimated at ($4.05 + $8.43)/2 or $6.24/lb foratotal of $15.2 x 106 . Dewar Since the winding and superstructure envelope were independent of current level, the dewar size also remained constant. The dewar material in the reference was primarily aluminum with a total weight of 3.84 X 105 kg. This was design combined with an estimated unit cost of $2.75/lb to yield adewar cost of $2.51 x 106 independent of current level. eJ POWER SUPPLY SUBSYSTEM This subsystemwas assumed to consist of (1) the basicpower supply, (2) the protective circuits, controls, and interlocks, (3) the bus work, and (4) the vaporcooled leads. The characteristics of each were considered, cost estimates developed, and a single cost factor generated for the subsystem as a function of current level. The peak power required to charge the magnet is directly proportional to the energy stored and inversely proportional to the charge time. The baseline system stores about 6.7 x 109 J, hence the peakpower required would be approximately 1 MW for charge times in the 4-16 hr range. Information from manufacturers indicated that 50-kA, 5-V solid state units bad been built for about $80 x 103 and that a 250-kA unit would cost an estimated $200 x 103 • Other units, with current ratings between 50 and 250 kA, were expected to scale roughly linearly with current. At these high current levels it was indicated that cost rose slowly with voltage; for example, a 2-V, 250-kA supply was estimated at only 10% more than a 1-V supply at the same current. To allow some ftexibility in voltage and power level selection, the estimated cost rangewas increased by 15% at the low end and 50% at the high end for the purposes of this study. The primary protection for a superconducting magnet system of this type is a discharge circuit consisting of a resistor across the coil terminals or a set of resistors Impact of High-Current Operation on tbe Cost of Soperconduding Magnet Systems 15 across magnet sections. Because of the high degree of stability, the only likely manner for a normal region to propagate is if a section of the magnet is not immersed in liquid helium as would be the case if there were an interruption in liquid helium supply to the main cryostat with a subsequent decrease in liquid Ievel. Voltage and temperature transients during quench were performed for selected cases at several current Ievels with imposed Iimits on the fraction of winding allowed to go normal. Terminal voltages for discharge were limited to 2 kV and, in all cases, the internal voltage and temperature Ievels were reasonable; hence, it was concluded that standard methods of coil protection involving use of single or multiple resistors across coil sections would be adequate. In addition, estimates were made for the size of bus work and vapor cooled Ieads (note that the latter will represent a significant current-level-dependent heat Ioad on the cryogenic system). These were sized at a current density Ievel which would allow operation without water cooling in the bus work or vapor cooling in the power Ieads for the duration of a discharge. Neither component, however, appeared tobe a major contributor to the cost of the system. When totaled, the estimated cost for the components of this subsystem ranged from about $2.1 x 105 at 10 kA to about $5.9 x 105 at 250 kA. CRYOGENIC SUBSYSTEM The refrigerator /liquefier requirements were estimated on the basis of heat Ioads attributed to dewar losses, vapor-cooled Iead requirements, and Iosses due to joule heating in joints between superconductors. The latter component is a major contribution with a high Ievel of uncertainty owing to the Iack of data on large-scale conductors. The required refrigeration and input power required were then estimated and are sumrnarized in Table I. The cost for the system was found using a and updated with a cost escalation cost/power relationship given by Strobridge factor correspondin} to 8% per year. Results ranged from $4.64 x 105 at the 10-kA Ievel to $1.53 x 10 at the 250-kA Ievel. e1 Table I. Requirements for BaseUne System* kA Lead requirement, liters/hr Number of joints Joint losses,t liters/hr Total helium, liters/hr RefrigeratorI liquefier efficiency, % ofCarnot Electrical power input,:t: 10 25 50 100 150 200 250 30 75 150 300 450 600 750 189 189 189 114 94 95 113 19 33 82 176 313 548 1004 152 211 335 579 866 1256 1856 14 15 17 18 19 19 20 231 292 376 579 772 1024 1264 Operating current, * Dewar Iosses for the baseline were 103 liters/hr. kW t Assumes superconductor joint resistance = 10-lO n. :t: Input power is based on assuming that the refrigerator /liquefier operates as liquefier for dewar and Iead Iosses (enthalpy of boil-ofl gasnot available to refrigerator) and as a refrigerator for joint losses. 16 R. J. Thome, R. D. Plllsbury, H. R. Sepl, lllld 8. 0. Pedenon FABRICATION AND ASSEMBLY The fabrication cost elements which are sensitive to current Ievel are those Operations associated with winding component preparation, winding, and assembly to the substructure. The cost elements which are relatively insensitive to current Ievel are those associated with assembly to superstructure, dewar, and the support systems. A simplified manufacturing blockdiagram is shown in Fig. 2. The operations up through "Oamp turn in Place" are current-level dependent and were divided into 12 fabrication tasks. Foreach task, a Iabor and tooling cost were generatedas a function of current Ievel based on (1) number of lengths of stabilizer received at the fabrication facility (= number of stabilizer joints), (2) number of lengths of stainless steel channel, (3) length of stabilizer, (4) lengths of stainless steel channels, (5) number of turns in saddle coils/racetrack coils, (6) number of lengths of superconductor received at fabrication facility, and (7) length of superconductor. Labor hours were estimated for each of the 12 tasks, engineering supervision assumed as equivalent to 15% of the technician Iabor total. lt was then assumed that technician Iabor would be billed at a rate of $15 per hour and that engineering Iabor would be billed at $35 per hour. The costs for these operations decreased from $19.5 x 106 to $7.32 x 106 from the 10-kA to the 250-kA Ievel with the greatest variation at the low current Ievels. The remaining Operations in Fig. 2 are independent of current Ievel. Labor estimates were generated and priced at $15 or $20 per hour depending on the task involved. Supervision was again assumed at 15% of the totallabor requirement and priced at $35 per hour. The total cost for these operations was estimated at $5.91 X 106 • INGOMING INSPEGTION OF STABILIZER, GHANNEL OTHER GOIL MATERIALS INGOMING INSPEGTION OF SUPERGONOUGTOR a ASSEMBLE STABILIZER SS GHANNEL PREFORM TURN SEGMENT; OR MAGHINE, THEN ASSEMBLE PLAGE TURN SEGMENT IN POSITION a BRAZE STABILIZER GROUNO WRAP a GURE INSULATION GLAMP TURN IN PLAGE a a ASSEMBLE SUPERGONOUGTOR TURN SEGMENT; SOLDER a a a APPLY SS STRIP SPIRAL WRAP TURN SEGMENT WITH INSULATION ASSEMBLE TO SUPPORT SYSTEMS Fig. 2. Manufacturing blockdiagram for coil assembly using separate substrate conductor. Impact of High-Corrent Operation on the Cost of Soperc:ondocting Magnet Systems 17 COST INTEGRATION AND CONCLUSIONS A total cost would have to be viewed with extreme caution since costs were estimated primarily on past experience and general discussion with selected potential vendors. Quotation drawings and specifications were not prepared and Submitted for vendor evaluation since this was beyond the scope of this effort. Furthermore, no allowance was made for design, quality assurance, facility preparation, or administrative expenses. However, the relative costs indicate trends and the exercise shows which factors are current-level dependent. The relative total cost is shown in Fig. 3 and the relationship between cost components is shown in Fig. 4. The results indicate that (1) the component costs for conductor, substructure, power supply subsystem, and refrigerator/liquefier subsystem, generally increase with current level; (2) the cost for fabrication of windings and substructure generally decreases with current level; (3) the combined effects of items (1) and (2) is to yield a relatively flat minimum in the total system cost in the vicinity of 100 kA. Alternate cases were also considered in which the allowable heat flux from the conductor surface was varied between 0.3 and 0.9 W /cm 2 • The net change in the bottom line total cost was a few percent. There are several factors inherent in the study which tend to favor the higher current levels. For example, the design assumes each conductor to be surrounded by a substructural channel. Although this is necessary at the high current levels to satisfy the local bearing pressure constraint assumed for the conductor, it is not necessary at the lower levels where several conductors might be housed in each substructural module. This approach would tend to make winding fabrication less costly at lower current levels. The study also assumed that the amp-meter requirement and winding envelope remained independent of current level. In fact, the overall current density would probably decrease somewhat as current level increased, therefore leading to a greater conductor requirement and a larger winding with a need for an increase in superstructure and dewar size. 1.2 1-- <n 0 u 0.8 ..J ""0 1-- 1-- "'> >= ""..J "' Q: Fig. 3. Relative total system cost vs. Operating current Ievel. 0.4 0 100 CURRENT, kA 200 18 R. J. Thome, R. D. Pilllbary, H. R. Sept, IIDd 8. 0. Pedenon 10° Tolal Cosl $ ,_"' z w z 0 a. c A ~ H I G 10 - 1 ::e B 0 <.> ...0 0 ,_ F "'<.>0 w > ;:::: * E 10-2 _, ~ w a: 8 • Substruclure C • Superstructure 0 " Oewar E • Power Su pply Subsystem 0 F • CryoQenic Subsystem G " Mlsceltoneous Componenls and Shlpptn-., H • Windl n9 ond Subtlructure Assembly I " Assombly to Superstructure, Oewar ond Support S ystems 100 200 CURRENT, kA Fig. 4. Variation of component costs; total cost for a given current Ievel is 1.0 (exclusive of design, quality assurance, facility preparation, or administrative costs). In addition, the lifetime cost of power to operate the cryogenic and power supply subsystems would tend to shift the minimum to lower Ievels if this operating cost were added to the capital cost of the magnet system. As a result of the above, the relative total cost curve would be expected to show somewhat less variation at the low-current end and more variation at the high current end with a slight shift of the minimum to lower Ievels. Though the effects must be evaluated in more detail and are dependent on the scale of the magnet, the implication is that future reference designs for large-scale MHD magnets should consider Operating current Ievels somewhat higher than the 10 to 20 kA typically used to date. REFERENCES 1. R. J. Thome, J. W. Ayers, T. M. Hrycaj, and R. D . Pillsbury, "Design of Superconducting Magnets for Magnetohydrodynamic (MHD) Applications," Final Report ERDA Contract No. E (45-18)-2217, June 1977. 2. R. J. Thome, R. D . Pillsbury, J. W. Ayers, and T. M. Hrycaj, IEEE Trans. Magn. MaJ-15(1) : 306 (1979). 3. T. R. Strobridge, NBS Tech. Note 655 (June 1974). A-3 FINAL DESIGN OF A SUPERCONDUCTING MHD MAGNE TFORmE COAL-FIRED FLOW FACILITY AT THE UNIVERSITY OF TENNESSEE SPACE INSTITUTE* S.-T. Wang, L. R. Turner, L. Genens, W. Pelczarski, J. Hoffman, J. Gonczy, H. Ludwig, R. C. Niemann, K. F. Mataya, and E. Kraft Argonne National Labaratory Argonne, Illinois and W. Young University of Wisconsin, Madison, Wisconsin INTRODUCf iON The superconducting magnet system (SCMS) consists of a superconducting magnet, magnet cryostat, a helium refrigerator/liquefier facility, a helium gashandling system, apparatus for cryogenic transfer and storage, a magnet power supply, an integrated instrumentation and control system including a computer for magnet operation, data acquisition, system status and diagnosis, and magnet protection. The complete systemwill be tested at Argonne and installed at the Coal-Fired Flow Facility (CFFF) at the University of TennesseeSpac e Institute (UTSI) in 1981. The cryogenic aspects and magnet safety are described and analyzed elsewhere C' 2 ]. The MHD warm bore, the magnet cryostat Iayout, and the axial field profile, are shown in Fig. 1. Significant magnet system parameters are listed in Table I. It is recognized that the UTSI SCMS must serve as a model with the design scalable to future large MHD magnets as the Engineering Test Facility (ETF) andin the full-size base Ioad systems. The successful experiences gained in designing and fabricating the UTSI SCMS will definitely advance the state-of-the-art of large MHD magnets and will provide a significant technological data base for future large MHD superconducting magnets. For a given conductor design, cooling provisions, and given coil structure, one can compute the minimum propagating energy and perform verification experiments. Nevertheless, because of the complex mechanical nature of a given coil and lacking acceptable criteria for the size of the disturbance against which the conductor should be stable, the UTSI SCMS conductor is designed to be absolutely or unconditionally cryostable. Despite the absolute stability, an achieved average * Work supported by the U. S. Department of Energy. 19 20 1 r-- S.-T. Wang etal. ~00 cm 6~HCTIVE FIELO ~ 6 FIELD TESLA 4.8 T 4.8 1 IN 3 AX I AL POSITION , cm -WARM B<JIE 108.8 - - t -- -300 cm HF. FIELD 80cm DI A. cm - aA. •~"--- ---'-- 1-- - -- - -- - - 1 640 cm - - - - -- - - -- -- - -l Fig. 1. MHD warm bore, cryostat Iayout, and field profile. current density of about 2000 A/cm 2 over the coil cross section is adequate so that the economy of the total ampere-turns needed is quite good. The coil configuration is the U . S. SCMS type circular saddle eJ. The coil will be assembled on the magnet bore tube with spiral banding. The bore tube will be about 6.3 cm thick (in the thickest section) and the banding will be strong enough for coil assembly but too weak to transmit the 30,180-kgr/cm maximum burst force to the bore tube. Fifteen ring girders will be used as the superstructure to contain the force. The decentering force of about 0.2177 x 106 kgt will be taken up by end ftanges and bore tube. COIL CONFIGURATION S As shown in Fig. 2, the magnet is circular saddle wound, with 14 layers. The outer 13 layers are each 4.88 m long; the inner layer consists of two subcoils: 1A, 2.04 m long; and 1B, 2.87 m long. This coil spacing provides the required field Final Design of a Superconducting MHD Magnet 21 Table I. Magnet System Parameters for CFFF SCMS MHD warm working bore size MHD on-axis field MHD etfective magnetic length Field ripple, uniformity and field tapper Ripple along field taper Cross-sectional field uniformity Field taper Magnetic coil characteristics Coil winding bore (minimum) Bore tube thickness (maximum) Winding type Operation current Peak field Winding build Average current density Cryostability Stored energy lnductance Max. Iransverse force Coil force support Support against external forces during operating and shipping Cryostat 4.2 K cold-mass weight 4.2 K cold-mass dimensions Liquid helium inventory Heat leak to LHe 4 Vacuum vessel, support and LN 2 shield weight Vacuum vessel dimension Total magnet weight Refrigerator /liquefier facility lnlet 80 cm, end of etfective field 100 cm Inlet at 4.8 T, peak at 6 T, outlet at 4.8 T 3.0m ±2.5% of on axis field ±5% of on axis field 0.2T/m 119cm -6.3cm Circular saddle with intermediate crossovers 3675 A 6.9T 0.53m -2890 A/cm 2 in copper; -2000 A/cm 2 in winding 0.142 W/cm 2 -168MJ 21H 30,180 kgdcm Partial support from bore tube and major support from external ring girders 3 g vertically down, 1 g vertically up, ±2 g axial, and ±1 g lateral 1.32 X 10 5 kg -3.16-m diam x 5.05 m long 8962liters 14.0 watts (=20 liters/hr) -4.06 X 104 kg -3.6-m diam x 6.4 m long -1.73X10 5 kg Closed-cycle refrigeration capable of system cooldown in 4 weeks and liquid helium production of 50 liters/hr. Other components are liquid helium storage dewars, helium gas recovery. profile. This coil cross section is chosen so that it is efficient in producing magnetic field and keeps the azimuthal force buildup in each layer to an acceptable Ievel. Layers two through nine all subtend an angle of 53°, ensuring that the radial force from each layer is supported by the next. CONDUCTOR DESIGN To reduce conductor costs, three grades of conductor, A, B, and C, designed for fields up to 7 .5, 6.5, and 4.5 T, respectively, will be used. The critical current is 4500 A at 4.2 K. Grade A will be used at fields up to the peak field; grade B, up to 6.0 T; and grade C, up to 4.0 T. Layer 1, containing only 42 turns, will consist of grade A conductor. Layers 2-4 will have both grade A and B conductors. Layers 5-8 will have all three grades; and the remaining six layers will have both grade B and C conductors. For the inner layers, the field where the layer begins to turn over at the high-field end determines the turn at which grades must be changed; for the outer layers, it is the field at the high-field end. .Z2 S.-T. Wang et al. SECTHll>l A-A SADDLE COILS BORE TUBE LOW FIELD END HIGH FIELD END TOP VIEW _ ::l, f---- - - - - 4 . 8 8 1 1 -- -- -----t Jlm 1_~~ RAD. .S91m lA RAD. _I.B - - - - - t j . SECTION B-B Fig. 2. Coil configuration of circular saddle coil. As shown in Fig. 3, the selected conductor will be 3.10 cm high with conductor thickness varying according to the field grades. Present design plans call for the superconducting insert to be a soldered cable of multiple wires of Nb-48% Ti superconducting composites transposed around a copper strip. The stabilizer will be an integrated piece with a longitudinal groove. The superconducting cable is to be bonded by 50-50 noncorrosive soft solder to the longitudinal groove of OFHC copper stabilizer. The depth of the longitudinal groove is such that the surface of the superconducting cable strips will be 0.76 mm below the front face of the copper stabilizer. COIL STRUCI'URE The turn-to-turn and layer-to-layer winding detail are shown in Fig. 4. The turn-to-turn insulation will be provided by a pultruded fiberglass strip with keystoned cross section. The keystoned feature eliminates the tedious Iabor of inserting angle-correcting wedges during coil winding. The pultruded fiberglass strip will be fashioned in a fishbone-like pattern so that vapor-locking will be avoided and that about 30% of the broad face will be cooled. The conductor designed with this face cooling and 50% edge cooling will be cryostable with a steady-state heat transfer of 23 Final Design of a Superconducting MHD Magnet TWENTY 1.02 mm TWENTY 0.89 mm DIA. SC WIRES-TWIST DIA. SC WIRES-TWIST FllliiENT Oll . ' PITCH ' STR lND DIA. a 10»... FILAMENT STR lND PITCH ' 0.2 TURN/tM 10 ... 0.2 TURM/tM lMNEALED COPPER AHNEllED COP PE A SEVENTEEN 0.79 mm DIA, SC WIRES-TWIST FllliiENT Oll,' 60..._ STRAND PITCH ' 0.2 TURN/tll AMM Ell ED COPPER Fig. 3. CFFF SCMS conductor. PARTIAL VIEW OF SADDLE COI L { NOT TO SCALE) PULTRUDED FISERGL ASS BANOING- SPIRAL WO UNO SUPERCONDUCTOR KEYSTONE·SHAPED PULTRUDED FISERGLASS SUPERCONOUCTOR ( 4.7 mm • 31 mm ) END TURN REGION STRAIGHT REGION TURN·TO·TURN STRUCTURE LAYE R· TO·LAYER STRUCTURE Fig. 4. Coil structure of turn-to-turn and layer-to-layer. 24 S.-T. Wang etal. about 0.14 W /cm 2 to the liquid helium. This low heat ftux assures the unconditional stability for the conductor. The CFFF SCMS is to consist of 14 coillayers. Except for coils in layer 1 which will be banded to the bore tube by 321 stainless steel, each coillayer, having two circular saddle coils, will be assembled onto the bore tube by pultruded fiberglass banding of 1.85-cm width by 0.71-cm thickness. The pitch length is 3.7 cm for the straight coil region and 2. 78 cm at the coil end regions. The pitch length is reduced at the coil end region because the conductors in the coil end are nearly parallel to the spiral banding at various locations. Thus about 50% of the conductor edges will be cooled in the straight coil region while about 30% of the conductor edges will be cooled in the coil end region. A micarta coil form, similar to the one used in the U. S. SCMS, will be employed to wind the CFFF SCMS saddle coil. The coil form will be fabricated from 3.084-cm-thick linen base phenolic Iaminate. The form is tobe slotted to allow it to conform to the cylindrical saddle shape. The slots will be filled with glass fillers and shell828-Versamid 140 epoxy to sustain the curved shape. It is planned to insulate the bore tube with insulating sleeves which are constructed of a wet-wound fiberglass cylinder of about 0.635-cm thickness. The cylinder will be machined with cooling grooves 1.428 cm wide by 0.476 cm in depth. A layer of 0.635-cm-thick G-10 coil blanket is tobe placed over the insulating sleeve cylinder to provide both 50% edge cooling and coil support at the coil end region. Similar coil blankets of 0.655-cm thickness are used to insulate the stainless steel banding by placing the coil blankets ahove and beneath the stainless steel banding of coillayer 1. CRYOSTABILITY Calculated recovery currents for the three grades of conductor are 3780 A at 7 T for grade A, 3860 A at 6 T for grade B, and 4300 A at 4 T for grade C. The design calculation is based on a heat transfer of 0.15 W /cm 2 • All such recovery currents are above the design operating current of 3675 A. Specialattention must be paid to the cryostability of the end regions, for three reasons: (1) Peak fields in Band C grade (not A) conductors occur at these regions; (2) there is less edge cooling, because the banding pitch is tighter; and (3) there may he lengths of conductor, up to 33 cm (13 in.) long, covered by the handing in a region where the handing and conductor run parallel. ELECTROMAGNETIC FORCES AND PRESSURE The conductor in the high-field (6.0 T) cross section A-A of Fig. 2 exerts a hurst force of 30,180 kg,/cm outward on the girders; a clamping force of 24,287 kg,/cm pushes the two halves of the magnet together. Figure 5 shows the radial magnetic pressure due to the radial and azimuthal components of force for each 5° azimuthal increment, totaled over all 14 layers to give the net pressure and also totaled to the layer which gives the peak pressure. The radial pressure shown in Fig. 5 includes the radial pressure due to the azimuthal forces; the comhined pressure gives the hurst force of 30,180 kgt/cm. Figure 5 also shows the accumulated pressure for each layer, calculated in the high-field cross section. The nurober of turns per layer were chosen to Iimit this average pressure to prevent yielding in the copper stahilizer of the conductor. The coil ends contrihute axial forces as weil as hurst and clamping force. Final Design of a Superconducting MHD Magnet 15 40 ..,. 30 PEAK RADIAL PRESSURE AMONG ALL LAYERS ( BOTTOM SC ALE) :::.J! w a:: => "'"' w a:: 20 <>. RAOIAL PRESSURE LAYER 14 ( BOTTOM 10 SCALE ) 0 10 20 30 40 50 55' DEGREES GAP . Fig. 5. Magnetic forces and pressure . The decentering forces on each coil are especially bothersome for a dipole magnet with a tapered field. These forces result from the asymmetric field distribution. Each layer experiences a different field at the two ends, and consequently different and noncancelling forces. The 219,500 k&r is distributed nearly uniformly over the outer 13 layers. 26 S.-T. Wang etal. STRUCfURAL SUPPORT AND STRESS ANALYSIS The bore tube provides some structural support for the cold mass and magnetic forces. The assembly of two end ftanges and a cylindrical section will weigh 13,152 kg. The bore tube details and dimensions are shown in Fig. 6. The derentering magnetic force is supported by the ftange at the high-field end and by a step in the tube. In addition to the magnet forces, the bore tube assembly is designed for a 3.4-atm (50-psi) pressure in the helium vessel. The end ftanges are annular plates of 2.34-m diameter supported at the inner boundary by the bore tube. The outer boundary attaches to the outer helium shell. Bending moments transferred by the end ftanges to the bore tube are supported by gussets on the tube section outboard to the ftanges. A total of 32 gussets (including four supportlink gussets) are provided on the high-field ftange and 16 (including four support link gussets) are provided on the low-field ftange. All gussets have a minimum thickness of 2.54 cm except the 10.16cm thick gussets at the cold-mass support link locations. The ftanges will be cast from ASTM A 351-7 5 grade CF8M modified stainless steel. This material contains a maximum of 0.04% carbon and is the cast alloy of wrought 316-L. Cast stainless properlies and soundness are enhanced with an increase of ferrite. Since the increased permeability associated with ferriteisnot detrimental to the design, ferritein the range of 5-15% has been specified. The yield stress of the ftange material will be at least 517 MNI m2 (7 5 kpsi) at 4.2 K. The maximum calculated stress 339 MN/m 2 occurs at the inner corner SUPPORT LINK LOAD UNE A G 8 I D II K II - 339 •250 88 1~0 33 1~7 29,& I~ 6Z IIAGNETIC PLUS IIEUUII LOAOING PRESSURE ~ 90"~ Fig. 6. Bore tube, end ftange, and stresses. Final Design of a Superconducting MHD Magnet 27 where the high-field end ftange joins the cylindrical tube. To minimize the ftange momentstransferred to the bore tube, a 22.86-cm (9-in.) stub of tubewill be cast as an integral part of the ftange. This positions the weld in a location where the moments have essentially died out. Fabrication of the tubular section of the assembly will be from rolled and welded annealed, and pickled, corrosion-resisting chromium nickel steel plate type 316-L, as per ASTM A240-75. The 4.547-m-long tubewill be fabricated from four rolled sections, solution annealed prior to welding. This method of fabrication results in one longitudinal and five circumferential welds. The stress value at various high-stress points are indicated in Fig. 6. The 11 main girder rings are 30.48 cm wide, spaced 2.54 cm apart, and symmetrically located over the 300-cm effective field region. Each ring consists of two 130° cast stainless steel segments pinned to six (three per side) aluminum tie plates as shown in Fig. 7. The cross section of the main girder segments is an 1-beam with a solid center portion. Thesesegments will be cast fromASTM A 351-75 grade CF8M (type 316 stainless steel). Selection of aluminum tie bars in the main girder rings was made to minimize the weight and to assure that the girders will be tight against the fiberglass bands after cooldown. Three 5.08 x 20.32 x 165.6-cm-long plates per side fabricated from 2219-T87 aluminum have been specified to meet the need of good ductility and high yield strength. The pin selected for the finger arrangement shown in Fig. 7 is 6.67 cm in diameter and is solution treated and aged A-286 stainless steel with an estimated yield of 896 MN/m 2 (130,000 psi) at 4.2 K. The conductors at each coil end will be supported by two end ring girders made of 90° segments of 2219 T87 aluminum alloy. The girder located closest to the high-field plane will have a radial thickness of 30.48 cm whereas the outer girder will be 20.32 cm thick. The segmentswill be cut from a 5.08-cm-thick plate stock. The maximum tensile stress in the body of the 20.48-cm ring girder is 193 MN/m2 • The maximum tensile stress in the girder at the pin holes is 234.4 MN/m 2 • POWER SUPPLY, INSTRUMENTATION, AND MAGNET PROTECTION The power supply will have an output characteristic of 200 V and 5000 A. A 0.05-0 dump resistor ~submerged in a water tank, will be used to safelydischarge the magnet in the event of an emergency. Design considerations indicate that the magnet can be energized in 70 min and discharged in 7 min. The maximum discharge terminal voltage will be 100 V with respect to the groundin-g center tap of the dump resistor. A total of 95 strain gauges are distributed among high-stress points. Thermocouples fabricated of copper-constantan, silicone diode, and Au-Fe are used within the coil winding, helium vessel, bore tube, ring girder, cold-mass support and nitrogen shield system. Potential taps will be needed for monitoring charging voltages and the normal transition of the conductor. Motion of the conductor will be monitored by an accelerometer and microphones in addition to the potential taps. Hall probes will be used to measure the MHD channel field and the peak field in the winding. Finally, the liquid helium Ievel and the liquid nitrogen Ievel will be measured by a NbTi probe and oxygen bellow assembly, respectively. MAIN RING GI RDER UNDERSIDE VIEW ) - -- ({!) - EJ1] .. END TOP \/ l EW 2219 - T87 ALUMINUM PLATE 6.67-cm- DIA. PIN E$cJ PLATE 2219·187 ALUMINUM ~5-cm-THK. ~,., _,,._ '" Fig. 7. Main ring girder and end ring girder. 2219-187 ALUMINUM TIE-PLATE 6.67-cm- DIA . PI N CAST SS RING GIRDER l fGRDOIIE PATTERN 2219-T87 ALUMINUM TIE- PLATE 6.67-cm- DIA. PI N Final Design of a Superconduding MHD Magnet 29 1. R. C. Niemann, S.-T. Wang, J. W. Dawson, L. Genens, R. P. Smith, L. R. Turner, J. D. Gonczy, J. Hoffman, K. F. Mataya, P. Smelser, P. C. Vander Arend, and S. Stoy, in Advances in Cryogenic Engineering, Val. 25, Plenum Press, New York (1980), p. 30. 2. L. R. Turner, S.-T. Wang, R. P. Smith, P. C. Vander Arend, and Y.-H. Hsu, in Advances in Cryogenic Engineering, Val. 25, Plenum Press, New York (1980), p. 39. 3. S.-T. Wang, R. C. Niemann, R. L. Kustom, P. Smelser, W. J. Pelczarski, L. R. Turner, E. W. Johanson, E. F. Kraft, S. H. Kim, J. D. Gonczy, H. F. Ludwig, K. F. Mataya, W. E. LaFave, F. J. Lawrentz, and F. P. Catania, in Advances in Cryogenic Engineering, Val. 23, Plenum Press, New York (1978), p. 17. DISCUSSION Question by E. Mullan, Westinghouse R & D Center: What is the expected power output from the MHD facility? Answer by author: The maximum power output from the MHD facility will be about 80 MW (thermal). Comment by T. Hrycaj, Magnetic Corporation of America: I do not understand how the fishboneIike turn-to-turn structure helps avoid vapor locking. In Fig. 4, the fishbone-like pattern appears to cut off any radial ventilation of gaseous helium between layers. Furthermore, the lower half of the fishbone-like pattern Iooks like it would actually trap rising helium bubbles. Answer by author: Vapor locking is avoided because the longitudinal groove of the conductor is recessed below the broad face of the conductor and also because the punched slots of the turn-to-turn insulator do expose the slots to the longitudinal groove of the conductor. A-4 CRYOGENIC ASPECfS OF THE UTSI-CFFF SUPERCONDUCfiNG DIPOLE MAGNET FOR MHD RESEARCH* R. C. Niemann, S.-T. Wang, J. W. Dawson, L. Genens, R. P. Smith, L. R. Turner, J. D. Gonczy, J. Hoftman, and K. F. Mataya Argonne National Laboratory Argonne, Illinois P. Smelser Independent Consultant Jefferson, Missouri and P. C. Vander Arend and S. Stoy Cryogenic Consultants, Inc. Allentown, Pennsylvania INTRODUCI'ION The Argonne National Laboratory, under the sponsorship of the Fossil Energy Division of the U. S. Department of Energy, has designed and is constructing a 6-T, 0.8-m minimum warm bore superconducting dipole magnet system for magnetohydrodynamic (MHD) research. The system will be installed and operated at the University of TennesseeSpace Institute (UTSI) Coal Fired Fuel Facility (CFFF). The systemwill consist of a coil assembly contained in a liquid helium cryostat, a helium refrigerator/liquefier slstem, and controls and instrumentation for cooldown and steady-state operation L1-4]. MAGNET CRYOSTAT The cryostat consists of a helium vessel surrounded by a liquid-nitrogen-cooled thermal shield contained in a vacuum vessel. The cryostat geometry is shown in Fig. 1. Steady-state heat fluxes to the helium vessel and the thermal shield have been minimized by the use of multilayer insulation along with low-heat-leak support * Work supported by the U. S. Department of Energy. 30 Cryogenic Aspects of the UTSI-CFFF Superconducting Dipole Magnet 31 Fig. 1. UTSI-CFFF MHD magnet geometry. structures. Instrumentation Ieads are sized to reduce the heat leak with intercepts employed to reduce conduction heat flux. Bafftes will be employed to reduce the radiation heat flux in penetrations. The steady-state heat flux to the helium vessel is 14.0 W or 19.6 liters/hr. The steady-state heat flux to the thermal shield is 179.6 W equivalent to 4.3 liters/hr boil-off. The cryostat design parameters are given in Table I with internal details shown in Fig. 2. WEIGHT SUPPORT RING SUPPORT TUBE LHe VESSEL VACUUII VESSEL ALUMINUM RING GIROERS COLO MASS SUPPORT WARM BORE PEOESTAL FLOOR PAO Fig. 2. UTSI- CFFF MHD magnet cryostat internal details. 31 R. C. Niemann et al. Table I. UTSI CFFF SCMS Cryostat Design ~arameters He4 Vessel Designpressures (gauge) lotemal 3.45 x 105 Pa (SO psi) Extemal1.03 x 105 Pa (15 psi) Weight Bore tube and end ftanges Coil assembly Ring girders Outer shell Total Uquid inventory Loads Radiation He4 vessel supports Leads (c:ontribution with excess cooling) 12R loss in conductor joints Penetrations, instrumentation Ieads, standoffs, etc. Total Tempersture Boil-off rate Material Thermal radiation shield Type Loads Radiation He 4 vessel support heat intercepts He4 vessel neck intercepts Standoffs, piping, etc. Total Tempersture Boil-off rate Weight Material Vacuum vessel Design pressures (gauge) Internal1.38 x 105 Pa (20 psi) Extemall.03 x 105 Pa (15 psi) Weight Shell Support structure Total Material: 304 ss 17,230kg 57,6SOkg 47,620 kg 7,070kg 129,570 kg (142.8 tons) 8,9601 94.1 m2 at 0.043 W/m 2 8 supports at 0.125 W /support 61 joints at 0.04 WI joint 4.5K 19.6liters/hr 316 ss (cast and wrought) = 4.17W = l.OOW = 1.48W = 2.44W = 4.89W 13.98 w Conduction cooled with Ioad into liquid nitrogen reservoir. Aluminized Mylar on both sides. 97.1 m2 at 0.645 W/m 2 8 supports at 3.02 W /support "'62.63W =24.16W 61.46W 31.41 w 180.00W 80K 4.3 liters/hr 2,180 kg (2.4 tons) 304 ss with copper tube liquid nitrogen tracing 17,510kg 20,950kg 38,460 kg (42.2 tons) The helium vessel, fabricated from 316 ss, is designed for the stresses of cooldown, steady-state operation, and fault conditions. Cooldown will be by forced convection of helium gas. Manifolding will provide uniform cooling of coil, ring girders, and helium vessel. Initial filling of the helium vessel will be through a tube to the bottom of the helium vessel. Refilling of the helium vessel will be through a tube which terminates in the upper regions of the helium vessel. The connections between 300- and 4.5-K environmentswill consist of two vertical penetrations at the end of the vessel. Two ports will provide redundant pressure relief. The helium vessel is suspended by a low-heat-leak support system employing tension-member-type supports. Epoxy fiberglass support members will be heat Cryogenic Aspects of tbe UfSI-CFFF Superconducting Dipole Magnet 33 intercepted at an intermediate point by connection to the thermal shield. The support system has four pairs of hangers affixed to each end of the helium vessel. The tension members will be preloaded to 50% of their maximum design Ioad to protect against possible impact loading. The helium vessel will be surrounded by a shield consisting of a liquid-nitrogencooled metallic envelope to intercept thermal radiation from 300 K. The shield has an average temperature of 80 K with a maximum temperature difference of 3 K. Considerations were made for eddy-current-induced forces in the shield associated with rapid discharge of the magnet. The shield incorporates a 304 ss sheet cooled by liquid nitrogen flowing in copper tubing. An externalliquid nitrogen supply reservoir sized to permit refilling a minimum of once every 48 hr will be employed. The space between the helium vessel and the thermal shield will contain ten layers of multilayer .insulation. The space between the thermal shield and the vacuum vessel will contain 50 layers of multilayer insulation. The shield will be instrumented with thermocouples to monitor performance and to provide for diagnostics during the cooldown and steady-state operation of the magnet. The assembly of the magnet cryostat will be completed by the installation of the vacuum vessel, designed as both a pressure vessel and a support structure. The 304 ss vessel will withstand the stresses of initial evacuation and purging, cooldown, steady-state operation, and fault conditions. Closure of the vacuum vessel will be at the helium vessel 300-4.5 K penetrations. Bellows will be employed to allow for differential thermal contraction without development of large stresses. THE CRYOGENIC SYSTEM The cryogenic system includes all elements for pumpdown and purge of the magnet cryostat, cooldown and filling from 300 K, and steady-state operation. The cryogenic system has been designed with emphasis on redundancy, high reliability, and safety. A schematic of the closed-loop helium flow system is shown in Fig. 3. A detailed arrangement of the cryogenic equipment is shown in Fig. 4. A helium refrigerator /liquefier will be employed to cool down the magnet from 300 to 20 K, to initially fill the helium vessel, and to provide liquid for steady-state operation. The refrigeration capacity is as required by the stipulated magnet cooldown time. The liquefaction capacity of 50 liters/hr will allow a redundancy factor of approximately 2.5 relative to the predicted liquid helium boil-off. Liquid nitrogen precooling will be used to aceeierate the early phases of cooldown. The cold box will employ turboexpanders. A helium compressor system provides process gas for the refrigerator/liquefier. The system can operate at approximately 50% rated capacity. The compressor system will be used to pump boil-off gas to storage. A 7500-liter liquid-nitrogen-shielded liquid helium storage dewar will be employed for filling the magnet, as a liquid reserve for times of equipment outages, and for storage of helium during magnet shutdown periods. Provisions will exist for the filling of the storage dewar for a liquid helium tanker which will be the primary source of liquid helium for the initial filling. Helium gas storage will consist of a single large storage tank having a capacity of 113,600 Iiters (30,000 gal) at a maximum storage pressure of 1.72 x 106 Pa (250 psig), i.e., approximately 2500 Iiters of equivalent liquid. A cryogenic console will contain control and diagnostic equipment for the operation of the cryogenic system. The console will be located at the refrigeration ~1.721106 Pa II g/ooc COMPRESSOR No. I u.__ ..."'"' ! 0 .."' w• Yn .. n i1.72x10 6 Pa 18 9/IOC No.2 COMPRESSOR ~vn !0 l LHo (7500 L ) DEWAR II 200W.4l 4.5 K 50 L/HR REFRIGERATOR/LIQUEFIER YVI"'Irn ,\.'r. ~-} He GAS RETURN 5'. COOLDOWN RET RN LHo MAGNET SUPERCONDUCTING- FILL LINE DEWAR SUPPL V f _?_ I fj'~ HEADER Fig. 3. UTSI-CFFF SCMS cryogenic schematic. ~ COOLDOWN ' I ( I I ~ I r- 8960 L LHo t 400L Jj t t I f FLOWMETE RS [][ ]!] ~ ..L.., ••• RESERVOIR ( '"• HEATER {p- 1- ro 1- SUPPLV Ho GAS HIGH• 000~'""' 1 ~ t • t • -· "r-W- 1900 L ~ r- '\..'r. LN 1 FLASH TANK r--1 1- @ 1.72 aiO' Po 113,600 L Ho GAS STORAGE MEDIUM- PRESSURE PRESSURIZATION 22,700 L U.T.S.I. LN 1 DEWAR WARM 9 1 LN 1 TO REFRIGERATOR f Tg -.::<2. I-- Cryogenic Aspec:ts of the UTSI-CFFF Superconducting Dipole Magnet LIQUEFIER/REFRIGERATOR 1 _ _,/V' '-I ,.- CRYOGENIC RELAY PANELS \ I 1\!IJj ~ \ l I \ SCMS CONTROL EQUIPMENT WITH CCIIPUTER -----. \ La ' coNTRoL LHe OEWAR 35 8 I I I BUILOING I I I I I ____ j HELIUM TRANSFER 'I LINES L- PIPING RACK ~*=t---rfl ~i TEST BUILOIN.G NORTH ( SCALE IN METERS ------'=-""""'""'=::----PO-WE_R_,SUPPLY, ETC. t 10 LsuPERCONDUCTING MAGNET Fig. 4. UTSI-CFFF SCMS cryogenic equipment arrangement. equipment for adjustments during cooldown and operation. Critical parameters required for steady-state operation will be remotely monitared and controlled from the CFFF control room. Liquidnitrogen will be supplied from the 22,700-liter (6000-gal) UTSI storage dewar and will be transferred to the magnet area in insulated lines. To assure the required liquid quality for operation of the refrigeration equipment, a 1900-liter (500-gal) flashtank will be installed in the cryogenics area and will serve as a centrat distribution point for liquid nitrogen to the cryostat and the refrigerator/liquefier. A utility vacuum systemwill be provided for the initial pumpdown and purge of the magnet cryostat, evacuation of transfer lines, refrigeration equipment, and other elements of the cryogenic system as required. A detailed listing of the major components of the cryogenic system is given in Table II. INSTRUMENTATION Instrumentation will be provided for the Operation of the magnet system which includes cryostat pumpdown and purge, cooldown, steady-state Operation, system diagnostics, and safety. The cryostat is monitared for lead gas flow, cold mass support forces, helium level, nitrogen Ievel, helium vessel pressure, insulating vacuum, and temperatures within the coil, the force containment structure, and the thermal radiation shield. The cryogenic system has been designed to permit remote monitaring of certain parameters. A remote instrument panel will be provided to monitor the critical operating conditions of the system. The panel includes readout of pressures, expander speeds, temperatures, and alarm signals for key components. Signals will be provided to the magnet system control computer for automated monitoring, analysis, and control. Remote control will be provided for J-T valve settings, cooldown bypass valve settings, compressor shutdown, emergency shutdown, etc. R. C. NieJaua et 111. 36 Table 0. UTSI CFFF SCMS Cryopnic Syste• Major Co•ponents Component Quantity Capacity 1 2 6.0-T field with 0.8-m minimum warm bore 50 liters/hr or 200 W at 4.5 K 0.04 kg/total (minimum) Magnet cryostat Refrigerator/liquefier cold box He4 oompressor liquid helium transfer lines Liquid helium storage dewar Helium gas storage tank Cryogenic oonsole Cryogenic rack liquid nitrogen storage dewar Liquid nitrogen flash tank Liquid nitrogen transfer lines Utility vacuum system Cold box to LHe4 dewar (ooaxial). liquid helium dewar to magnet cryostat cold box to magnet cryostat-2 1 1 1 1 1 1 From liquid nitrogen dewar to flash tank, to cold box, and to magnet cryostat 1 7500liters 113,600 Iiters (30,000 gal) at 1.72 x 106 Pa (250 psig) N/A N/A 22,700 Iiters (6000 gal) 1900 Iiters (500 gal) N/A OPERATION The cryostat will be cooled down with helium gas from the refrigerator/ 1\quefier. A full-ftow liquid nitrogen precooler will be included in the cold box. Flow to the cryostat is through a vacuum jacketed transfer line. The helium distribution piping is shown in Fig. 5. Gas is distributed to the bore tube and the ring girder areas through internal gas headers. Flow to the bore tube is through four ftow channels machined axially on the bore tube. Flow to the ring girders is through multiple headers located between the girders. Distribution of ftow between the bore tube and ring girders is accomplished with a valve located in the supply to the ring girder headers. The results of the cooldown calculations are shown in Fig. 6. This figure indicates the temperature vs. the relationship for the centrat cold bore tube, the middle layer coil (i.e., coil No. 7), the outer shell of the helium vessel, and a composite for the integrated magnet cold mass. The calculated cooldown time from 300 to 20 K is 24.9 days with a ftow of 0.04 kg/s of helium gas through the precooler. To evaluate potential thermal stress problems associated with temperature difterentials within the cold mass, a detailed cooldown analysiswas made. Because of the large size of the coil and the generally porous characteristics of the final assembly, laminar ftow exists with the resulting film coefficient for heat transfer being as low as 24 W /m2 K. However, the large surface area for heat transfer results in the cold helium gas quickly warming up to structure temperatures. By evaluating the energy balance and heat transfer relationships for each sector of the magnet, the temperature of each component vs. time was calculated. Once the cold mass reaches 20 K, cooldown will consist of liquid transfer from the storage dewar. The time from 20 to 4.5 K and filling with liquid helium has been calculated to be 53.6 hr, the total amount of liquid helium employed for this Operation being 12,366 Iiters. Cryogenic Aspects of the UTSI-CFFF Superconducting Dipole Magnet 37 BORE TUBE COOLING HEAOERS FILLE R WITH HO LES Fig. 5. UTSI-CFFF MHD magnet cryostat helium piping. Once filled, the Ievel will be maintained by periodic filling from the storage dewar. Because of the large liquid inventory above the coil assembly, i.e., 1550 Iiters, it will be possible to batch fill the magnet with intervals as long as 24 hr between transfers. During non-MHD Operating periods, the magnetcold mass will be kept at a temperature below 20 K to eliminate possible differential thermal motion which could require careful subsequent energization. 300 280 260 240 220 100 "'~~ ... ~ .... "'0.. ....::11 ~ 180 1~0 140 120 100 80 60 40 20 COOL DOWN TIME . OAYS Fig. 6. UTSI-CFFF SCMS cold mass temperature vs. time. 38 R. C. Niemann et al. FAULT CONDmONS AND RESPONSES The cryostat will be equipped with a vent system to handle major faults of the system leading to a massive heat flux to the liquid. The following two faults were considered: 1. The coil goes normal and the normal zone propagates. 2. Massive vacuum insulation failure occurs and the helium vessel becomes a huge cryopanel. Calculations are based on a case where the connection between the environment and the vacuum space are large enough so that heat transfer to the helium vessel becomes a limiting factor. Once the liquid outside the coil is warmed, fluid flow will start and the warmer liquid will enter the coil volume. This will initiate quenching of the magnet and additional fluid needs to be vented to maintain cryostat pressure at a safe Ievel. A vent system for the liquid helium cryostat has been designed for a maximum flowrateof 15.0 kg/s,amaximumpressurein the cryostatof 2.3 x 105 Pa (19.1 psig) and the temperature of the fluid in the vent line of 16 K. The vent system consists of a nominal 0.15-m(6-in.)-diameter pipe at both the 300- to 4.5-K penetrations. Cryogenic-system-related fault conditions have been considered with those factors having the ability to affect magnet system operation being pressure variations resulting from liquid transfer and refrigerator/liquefier malfunction. The effect of pressure variations is minimal owing to the large thermal mass afforded by the large liquid inventory of the cryostat. Refrigerator /liquefier malfunction should not directly affect the operation of the magnet since filling is from the storage dewar. Operations of the magnet with MHD channels should provide minimal perturbing influences. The warmborewill not be loaded by the channel so that direct vibrational coupling should not exist. The warm bore will be provided with an insulating Iiner capable of withstanding 10,000 V to ground. ACKNOWLEDGMENTS The authors wish to acknowledge the contributions of the following individuals to the design effort: J. Kotora, D. Gacek, E. Kraft, W. Sajdak of Argonne National Laboratory; R. Waiting, independent consultant; D. Krajcinovicof University of Dlinois Circle Campus; M. Hila! of the Michigan Technological University; and W. Young of the University of Wisconsin. HEFERENCES 1. S.-T. Wang, L. R. Turner, R. C. Niemann, L. Genens, and M. S. Srinivasan, "A Superconducting Dipole Magnet System for the MHD Facility at University of TennesseeSpace Institute," presented at 13th lntersociety Energy Conversion Engineering Conference, San Diego, California, August 20-25, 1978. 2. S.-T. Wang, R. C. Niemann, L. R. Turner, L. Genens, W. Pelczarski, J. Gonczy, J. Hoffman, Y.-C. Huang, N. Modjeski, and E. Kraft, IEEE 'li'ans. Magn. Mag-15(1):302 (1974). 3. S.-T. Wang, R. C. Niemann, L. R. Turner, L. Genens, W. Pelczarski, J. Gonczy, J. Hoffman, Y.-C. Huang, N. Modjeski, E. Kraft, K. Mataya, H. Ludwig, B. Phillips, J. Dawson, J. R. Purcell, W. Young, S. Stoy, and P. C. Vander Arend, "A Superconducting Dipole Magnet for the UTSI MHD Facility," presented at Superconducting MHD Magnet Design Conference, Cambridge, Massachusetts, October 18-19, 1978. 4. S.-T. Wang, L. Turner, R. Niemann, L. Genens, W. Pelczarski, J. Dawson, S. Kim, R. Smith, N. Kim, J. Gonczy, J. Purcell, J. Stekly, W. Young, J. Zar, P. Vander Arend, S. Stoy, and P. Smelser, "A Superconducting Dipole Magnet System for the CFFF MHD Facility at the University of Tennessee Space Institute," presented at 18th Symp. on Engineering Aspects of MHD, Butte, Montana, June 18-20, 1979. A-5 SAFETY ANALYSIS OF THE UTSI-CFFF SUPERCONDUCTING MAGNET* L. R. Turner, S.-T. Wang, and R. P. Smith Argonne National Laboratory Argonne, Illinois P. C. Vander Arend Cryogenic Consultants, Inc. Allentown, Pennsylvania and Y.-H. Hsu General Atomic Company San Diego, Califomia INTRODUCI10N Argonne National Labaratory is building a large superconducting dipole magnet [ 1] for MHD research at the University of TennesseeSpace Institute-coal Fired Flow Facility (UTSI-cFFF). In designing such a large (170-MJ stored energy) magnet, great attention must be devoted to the safety of the magnet and personnel. The conductor for the UTSI-cFFF magnet incorporates significant copper and contributes to the magnet stabilizer' which both insures its cryostability safety. This paper first presents the quench analysis, followed by the cryostat fault condition analysis. Two analyses of exposed turns follow: The first shows that gas cooling protects uncovered turns, and the second that the cryostat pressure relief system protects them. Finally the failure mode and safety analysis is presented. el MAGNET QUENCH ANALYSIS The behavior of the magnet in the event of a quench is modeled using the quench analysis program OUENCH [3]. The code calculates the longitudinal quench velocity for a conductor not in contact with helium coolant. For cases involving liquid helium, this velocity can be provided as a starting parameter. The transverse quench velocities can be provided from a calculation based on appropriate thermal properties. The quench volume of the coil grows in time as an ellipse whose axes are the * Work supported by tbe U. S. Department of Energy. 39 6 7 8 9 10 11 12 13 14 15 16 5 1 2 3 4 Case 108 109 105 99 112 118 118 125 156 224 242 246 266 302 325 236 Decay time constant, s 54 63 51 67 71 115 146 124 173 145 241 276 243 55 59 61 Peak temp., K 600 584 619 643 673 517 545 526 640 286 252 248 219 180 200 538 Peak inductive voltage, V 0.303 0.294 0.309 0.319 0.404 0.263 0.278 0.268 0.396 0.158 0.138 0.134 0.116 0.083 0.071 0.327 n Final resistance, 3.50 2.00 15.00 31.00 7.50 15.00 1.00 0.75 0.32 0.55 0.55 0.32 0.32 15.00 0.05 0.32 Longitudinal quench velocity, m/s -5% -5% -5% -5% -5% -5% -5% -5% -5% -1% -1% -1% -1% -0% -5% -1% Ratio of transverse quench velocity to longitudinal 0.73 0.71 0.75 0.76 1.00 0.69 0.68 0.65 1.00 0.40 0.37 0.35 0.31 0.24 0.19 1.00 Fraction energy in coil 0 0 0 0 0 20 0 0 0 0 24 0 24 0 0 0 %Helium Table I. Calculated Quench Results for Various Choices of Quench Velocity, Coil Protection, and Helium Cooling r I ;z: :< E. J. ~ I.. il J:1;1'.1 :- PI' I ~ ~ II PI' r ~ 41 Safety Alllllysis of the UTSI-CFFF Supercouducting Mapet 300 ...a:" ::> !C 200 ... ... a: 0.. :Ii 1- ...J ö c.J ..." 100 c( 0.. Fig. 1. Calculated coil peai( temperature for various quench conditions as a function of calculated current decay time. The numbers refer to the various cases listed in Table I. L-----~~----~L___--~~~ 100 200 300 TIME-CURRENT DECAY. s respective velocities multiplied by the elapsed time. The ohrnie heating generated in the winding structure as the stored energy of the field is dissipated can be calculated by the program, along with the time constant of the decay, the energy extracted by the protection resistor, the peak coil temperature, and relevant voltages. The conductor immersed in liquid helium is unconditionally stable, i.e., any normal zone in the conductor collapses. But, when there is no liquid helium in the winding voids, the normal zone will spread with a velocity of =750 cm/s. The cryostat fault condition analysis in the following section is based on this "no helium" calculation. Table I shows the peak temperature and peak inductive voltage for several sets of quench parameters. Without the protection circuitry described elsewhere, [ 1] the fraction of the stored energy deposited in the coil is 1.00. The vaporization of local liquid helium is not included explicitly in the OUENCH program. It can be modeled by increasing the enthalpy of the conductor. The last column of TableI refers to this additional enthalpy. Since the code makes no heat transfer calculations, the presence of liquid helium in the winding is modeled by a lower quench velocity. To see how important turn-to-turn and layer-to-layer heat transfer is, the program was rerun with the parameters which specify heat transfer reduced to values corresponding to a configuration with longitudinal propagation only. The table shows the peak coil temperature, peak inductive voltage, and decay time constant for each of the cases. Figure 1 shows the coil peak temperature, i.e., the temperature of the region where the quench initiated, as a function of the current decay time constant. CRYOSTAT FAULT CONDmON ANALYSIS The cryostat must be equipped with a vent system to handle major faults of the system which provide a massive heat ftux to the liquid helium in the cryostat. The following two conditions are considered. Magnet Quench The normal zone grows at a velocity of =750 cm/s for a coil in a helium gas atmosphere. In calculating the required rate of venting, it is assumed that there is 42 L. R. Turner, S.-T. Wang, R. P. Smith, P. C. Vander Arend, and Y.-H. Hsu liquid in the coil void fraction. This case appears to be very conservative because it is not real. Ha coil really quenches at this rate, liquid helium is not present and does not require to be expelled from the coil. But, if liquid helium were present, the normal zone would grow very slowly if at all. To be conservative, velocity along the conductor is set at 15 m/s (twice the value obtained above) with turn-to-tum and layer-to-layer velocities about 16 times lower. The cryostat contains a large amount of liquid helium in poor thermal contact with the coil itself, which does not participate directly in the transfer of heat from the coil. After the quench is initiated, boiling occurs inside the magnet structure, and the vapor generated displaces liquid helium out of the coil. The displaced liquid rises in the stack of the cryostat [4 ] and forces gas into the vapor retum line to the refrigerator; but the vapor line capacity is limited and pressure rises in the cryostat. The quench relief valve is set at 2.3 atm (19.1 psig), and approximately 0.5 s is required to expel enough liquid from the coil to compress the gas in the vapor volume of the stack to 2.3 atm. At that pressure there is no more boiling in the coil and heating of a single-phase fluid occurs, starting circulation of fluid through the coil. To maintain a constant pressure in the cryostat, the rate of venting must equal the increase in volume from the heating in the coil. If the fluid vented is at 2.3 atm and 15 K, the vent line required consists of a 6-in. (0.015-m) IPS pipe equipped with a relief valve capable of handling 9000 g/ s of helium vapor. To reduce pressure drop in the line, it is important that the entrance should be shaped to avoid the effects of a sharp-edged orifice. Massive Failure of Vacuum Insolation Consider the situation where the system experiences a massive vacuum failure to air. Boiling occurs on the inside of the cylindrical cryostat wall and the vapor generated displaces liquid into the stack [4 ]. Gas in the stack is compressed until the pressure reaches 2.3 atm. This takes a very short time because of the extremely high heat flux into the liquid. The quench relief valve [4 ] opens and maintains cryostat pressure at 2.3 atm. Boiling stops and the cryostat contains single-phase fluid. Fluid circulates in the space between cylindrical shell and girder rings Cl due to density differences, with velocities of the order of 40 cm/s. The outside wall of the cryostat is at approximately 70-7 5 K, in order for air to liquefy. Free access of air to the cryostat has been assumed. Heat flux through the cylindrical wall is of the order of 1.25-1.5 W I cm 2 • The temperature of the liquid rises, and the volume increases as determined by the temperature rise and mass flow rate along the wall. This increase in volume needs to be vented to maintain constant cryostat pressure. The required rate of venting is approximately 7200 g/s. The fluid temperature initially is approximately 16 K. Practically all of the heat added to the cryostat passes through the cylindrical shell. Rate of heat input to the vessel is on the order of 525 kW. The minimumrate of condensation of air (with free access to the ~ostat) is 1100 g/s. Even for air at sonic velocity, a minimumhole area of 60 cm2 (such as a 3-in. pipe) is needed to supply air at that rate. It appears that a massive vacuum break to airwill automatically be followed by a quench of the magnet, but the rate of mass expulsion of the warm fluid in contact with the normal coil is not as fast as for a quench without vacuum failure. The mass expulsion from the cryostat may be as high as double the rate calculated for quench alone. Finally, a vacuum failure of the liquid helium cryostat is much less serious than a vacuum failure to air because the helium gas transfers heat to the cryostat at a much Safety Analysis of the UTSI-CFFF Superconducting Magnet 43 60 " 50 .... <I 40 "'0::=> 0:: "'... :Ii "'.... " "'... <I 30 20 TI ME, s Fig. 2. Recovery of uncovered length following 1-kJ perturbation. lower rate than liquefying air. In fact, rupture of the helium vessel into a vacuum space open to airwill immediately reduce the rate of heat transfer from condensation because of a significant lowering of the partial pressure of air. ANALYSIS OF EXPOSED LENGm: EFFECf OF He GAS COOLING The analysis of an uncooled length of conductor, as might occur from the liquid Ievel falling below the top of the coils, is carried out with the BACKSAFE program using the back difference method, which permits the time increment to be changed during computation without mathematical instability. A heat balance is applied to each conductor element; the temperature changes under the combined effect of perturbing heat input, Joule heating, heat transfer to the coolant, conduction along the conductor, and turn-ta-turn conduction between conductors. The curve for convective vapor cooling of the uneavered length of conductor was obtained from the heat transfer coefficients calculated for the Rayleigh number, using the temperature-dependent properties of 1-atm helium vapor. Figure 2 shows the recovery of the conductor carrying 3675 A from a perturbation with a 1-kJ energy addition. The figure shows that recovery is identical whether the 1 kJ was initially distributed among 0.20 m of conductor at 72 K or 0.40 m at 57 K. If the conductor is cooled by liquid, it recovers in about 6 s. Fora 6-m one-half turn uneavered length, the temperature drops to 20-25 K in about 60s; subsequently, if the magnet is being dumped with a time constant of 420 s, temperature recovery is complete in 409 s. Whether or not the dump circuit acts, the magnet is not harmed by a 1-kJ heat input, even assuming the conservative cooling curve of 1-atm helium gas. ANALYSIS OF EXPOSED LENGm: EFFECf OF PRESSURE RELIEF SYSTEM This section shows that the CFFF magnet is self-protecting even if the dump resistor circuit fails while the helium level is low and a normal zone appears. 44 L. R. Turner, S.-T. Wang, R. P. Smith, P. C. Vander Arend, and Y.-H. Hsu The magnet has good thermal conductivity along the conductor (x direction), and poor thermal conductivities along the turn-to-turn direction (z direction) and the layer-to-layer direction (y direction); the latter two are about 2 x 10- 3 times the former. The conductor is unconditionally stable; consequently, any normal zone beginning above the liquid Ievel will not propagate below. For heat transport in a layer (x-z plane) parallel to the liquid helium surface, the contribution in the y direction is considered as a step function; a layer is counted as soon as a section of that layer emerges from the liquid. The temperature distribution along the y and z directions are assumed to be the same. Thus, the problern is that of a twodimensional heat transport with liquid helium cooling just below the liquid Ievel. The cooling by helium vapor is ignored. Throughout the analysis [5 ], the steady-state boiling liquid helium heat transfer is used. The latent heat of helium is adjusted to include the enthalpy needed for raising the localliquid helium from 4.2 K to the boiling temperature. The average temperature of gaseous helium is chosen as 5 K to obtain the pressure as a function of time. In the calculation [5 ], the normal region spreads out in the x, y, z directions and the dump resistor circuitry is inactive. The temperature distribution of the top layer, the heat generation rate, the pressure in the cryostat, and the liquid helium Ievel are all calculated. The results shown in Figs. 3 and 4 are for the CFFF conductor with an initial current of 3.675 kA and an initial vapor volume of 1200 Iiters at a pressure of 1 atm. The peak temperature is about the same for different initialliquid Ievels; typically, it rises up to 71 Kin 70s. Figure 3 shows the pressure and liquid Ievel as functions of time for different initialliquid Ievels. The buildup of pressure and the lowering of liquid Ievel are faster 2.5 2 .3 0 .10 E ...J 0 <..> ...._ 2 .1 0 ...0 Cl. E 1. 9 " "'~ "'"' "' "' 0.. ~ CAS E 2 , LIQ UID LEVEL 0 ...J "' CD "'<..>z 1.7 <t t- 1.5 CASE I' PRESSURE "'ä ...J "'> "' ...J 0 5 0 :::; 1.1 T I ME, s Fig. 3. Calculated liquid helium Ievel and pressure vs. time. Safety Analysis of tbe UTSI-CFFF Sopercondocting Mapet 45 3.6 PEAK TEMPERATURE 3.2 2.8 "" "'a::::> 2.4 ..... <( CURRENT a: "':I! "'..... "" "' 2.0 0.. <( 50 1.6 <( "" ..... z "'a:a: :::> <> 40 1.2 0.. 30 20 10 0 60 70 80 90 100 110 120 TIME, s Fig. 4. Peak temperature and current vs. time after all helium is removed. for the lower initial liquid Ievel because a larger amount of conductor rapidly generates and dissipates more heat into the liquid. Since the average temperature of the vapor will probably be higher than 5 K, the calculated pressure is somewhat underestimated. Also, the turn-to-turn cooling is neglected so the calculated conductor temperature is slightly higher than the actual temperature. For the worst case in which initially only part of the top layer is involved, it takes about 70s to build up pressure to 2.05 atm with the vapor temperature at 5 K and the peak temperature at 71 K. If all the helium is expelled, and there is no cooling at all afterward, the peak temperature and current as functions of time are as shown in Fig. 4. The peak temperature is only 89 K, and the current decreases to 2% of the initial operating current in 135 s. If the pressure relief valve is set at 3.0 atm, the peak temperature is about 110 K, still a safe value. FAlLURE MODE AND SAFETY ANALYSIS The possible failure modes, with their probabilities, associated hazards, and recommended actions are listed in Table II. Since the UTSI-CFFF magnet has an unconditionally stable coil, a trouble-free cryostat, a conservative power supply, and a carefully planned alarm, protection and diagnostic system, failure events associated with these items are estimated in a low probability category. On the other band, less critical item failures such as cryogenic piping rupture are estimated in a medium category, while a facility power failure is a likely event. The CFFF is in Seismic Zone No. 1, which means that earthquakes may cause minor darnage to structures with periods greater than 1 s and intensities that fall in the V and VI range on the M.M. scale. The probability of earthquake occurrence at UTSI is low. Nevertheless, the CFFF magnet and its associated equipment are designed so that they will remain in place and be functional following a major earthquake in Zone No. 1. Conductor going normal Loss of insulating vacuum Magnet coil open circuit Magnet/ cryostat Magnet/ cryostat Magnet/ cryostat Magnet/ cryostat Magnet/ cryostat Magnet/ cryostat Magnet/ cryostat 2 3 4 5 6 7 Over current or over field Current Iead too hot and open circuit Helium vessel relief valve fails to open during quench Liquid nitrogen fill valve fails to close Failure mode Items Systems components Low Low Low Low Low Low Low Estimated probability Arcing in coil, Ieads, power supply, dump switch Ring girder break Liquid nitrogen overflow Dump resistor dumps magnet energy Dump magnet energy; pressure relief valve open; helium vent to atmosphere (outside building); air may be liquefied Arcing in coil, Ieads, power supply, dump switch Burst disk will open Failure effects Electrocution if in contact with these systems Winding break apart Freeze 0-ring or flesh on contact Electrocution if in contact with these systems Explosion due to liquid oxygen Explosion due to liquid oxygen None Hazard description Table II. CFFF SCMS Fallure Mode and Safety Analysis Avoid 0-ring freeze/personnel contact Personnel must be isolated from these systems Provide escape measures Personnel must be isolated from these systems Provide escape measures Provide escape measures None Recommended action f ;:c ~ I r::r. = J. ~ i~ < ~ ~ J ~ !I' \G = ~ !"I 5" I I !I' r * Explosion or uncontrolled attraction by magnet Fallure to open during energy dump Not enough water in the tank during normal operation High-pressure gas bottle (He or N2) Dump resistor Dump resistor Power cables to magnet Facility power 11 12 13 14 15 Broken cable between dump resistor and power supply Power fails Rupture Cryogenic piping 10 High Low Low Low Low Medium Low Loss of insulating vacuum Liquid nitrogen ftash 9 Low Loss of insulating vacuum Liquid He storage dewar 8 Injury to personnel or magnet cryostat Release high-pressure gas or high-velocity impact Coil temperature will rise to 100K May open circuit unless dump switch is closed Freewheeling May arc in coil, Ieads, power supply, dump switch None May have some hazard as coil opens circuit unless dump switch is closed in time Electrocution if in contact with these systems None Injury to personnel Flying debris or released liquid Emergency relief valve open Explosion possible if liquid oxygen produced Explosion possible if liquid oxygen produced Emergency relief valve open Provide battery operated magnet current readout Personnel must be isolated from these systems Interlock dump switch operation to dump resistor water Ievel None Locate dewar in secured area and provide escape route Locate tank in secured area and provide escape routes Secure this piping and identify these lines with warning signs Secure and locate to remote area • i ~ II Ia Ig. .. f... ~ h ~ II Er s. ~ t'll 48 L. R. TIII'IIU, S.-T. WIIJII, R. P. Slllitla, P. C. Vander Arnd, and Y.-H. Hsu REFERENCES 1. S.-T. Wang, L. R. Turner, L. Genens, W. Pelczarski, J. Hoffman, J. Gonczy, H. Ludwig, R. C. 2. 3. 4. 5. Niemann, K. Mataya, E. Kraft, and W. Young, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 19. L. R. Turner, S.-T. Wang, and J. Harrang, IEEE Trans. Magn. Mag-15:371 (1979). M. N. Wilson, "Computer Simulation of the Quenching of a Superconducting Magnet," Rutherford Laboratory, RHEL/M/151 (1969). R. C. Niemann, S.-T. Wang, J. W. Dawson, L. Genens, R. P. Smith, L. R. Turner, J. D. Gonczy, J. Hoflman, K. F. Mataya, P. Smelser, P. C. Vander Arend, and S. Stoy, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 30. Argonne National Laboratory Superconducting Magnet Group, "Final Design of a Superconducting MHD Magnet for the CFFF-UTSI," Argonne National Laboratory, ANL-MHD-79-12, Vol. II (1979). A--6 ENGINEERING ASPECfS OF CRYOGENIC LASER-FUSION TARGETS D. L. Musinski, T. M. Henderson, R. J. Simms, and T. R. Pattinson KMS Fusion, Inc., Ann Arbor, Michigan and R. B. Jacobs R. B. Jacobs Associates, Inc. Boulder, Colorado INTRODUCI'ION For efficient burn of the deuterium-tritium fuel contained within a hollow, sphericallaser fusion target the fuel core must first be driven to a high density and then subsequently be elevated in temperature to initiate the reaction. To achieve high final fuel densities the internal pressure within the target must be overcome during the implosion. A cryogenic target, one in which the fuel is condensed as a liquid or solid layer on the inner surface of the spherical shell, may overcome the mechanisms [ 1] which can Iimit the final density. Fuel initially confined to the wall of the target cannot respond quickly enough upon absorption of energy to fill the interior volume of the target before the target implodes. At a given Ievel of Iaser power, a cryogenic liquid or solid layer target should compress to a higher fuel density and produce a higher yield than a target containing the same mass of fuel in the gaseous state [2 ]. The expected advantage of cryogenic targets can be illustrated by considering the specific example of a target filled with fuel at 100 atm. If irradiated at room temperature, the entire amount of fuel is available to absorb energy and impede the target implosion. However, if irradiated as a liquid-layer target at 21 K (approximately 1 K above the point where the fuel starts to solidify) only a small fraction of the fuel is in the gaseous state, contained within the inner boundary of the liquid layer. Raoult's law applied to a specific fuel fill of 8.4% HD, 9.3% HT, 19.7% 0 2 , 46.2% DT and 17.0% T 2 [ 3 ] indicates that the pressure within the interior volume of this liquid-layer target is 293 torr. At room temperature, the pressure of this amount of gaseous fuel would be 5.5 atm. Thus, only 5.5% ofthe total amount of fuel is available to absorb energy and impede the implosion if the target is irradiated as a liquid layer target. If the same target is irradiated at 15 K as a solid-fuel-layer target, the pressure within the interior of the target is only 11 torr. This corresponds to less than 0.02% of the total amount of fuel. Further lowering of the target temperature 49 50 D. L. Musillski. T. M. Heodenoo, R. J. Simms, T. R. Pattiosoo, md R. B. Jacobs Iikewise lowers the vapor pressure and hence lowers the amount of fuel remaining in the gaseous phase. A solid-fuel-layer target irradiated at the minimum temperature attainable should, therefore, yield optimum results. An extensive program has been undertaken to test these expectations experimentally. This paper discusses an essential element of the program-the engineering technology developed to integrate cryogenic systems into actual fusion target chambers. LIQUID-LAYER AND NONUNIFORM SOLID-LAYER TARGETS From an engineering point of view, the most direct way to produce a cryogenic target is by use of the technique referred to as point-contact conduction cooling [4 ] (as illustrated in Fig. 1). The target is bonded to the end of a 17-JLm-diameter copper fiber which is one element of a direct conduction path from the target to a source of refrigeration at Iiquid-helium temperatures. This technique is chiefly applicable to liquid-layer targets, i.e., targets in which the fuel is condensed as a liquid layer that completely covers the inside surface of the containing shell. Once the liquid layer has been established and stabilized by setting and stabilizing the temperature of the cryostat, target manipulation and alignment within the experimental target chamber proceed in the same manner as for a gaseous target at room temperature. Thus, this 7 ] to be conducted with minimum distechnique allows cryogenic experiments ruption of the normal routine of the target chamber. During the experimental sturlies on liquid-layer targets it became apparent that absorption of room temperature radiation dominates the dynamics [8 ] of the fuel layer. The liquid-layer configuration observed is the system's response to the absorbed radiative energy. The most direct engineering way to reduce this energy and its subsequent effect on the fuel is by the use of a radiation shield. Figure 2 shows, in schematic form, the integration of a simple cryogenic shield (or shroud) in the experimental chamber. A modular increase in the system's e- r· t5cm 8.5cm 1 l ~COLD ;t:-rr ~ENSK)N TIP 2cm INDIUM CRUSH WUHER Fig. 1. Point-contact conduction cooling cryostat. Engineering Aspects of Cryogenic Laser-Fusion Targets 51 StiROUO PISTOM Fig. 2. Integration of simple cryogenic radiation shield in target chamber. complexity, this shield allows a nonuniform solid fuellayer tobe formed within the target. A solid layer of DT fuel that completely covers the inside surface of the target (Fig. 3) can be formed by starting with a liquid layer that is protected under the radiation shield [4 ] and quickly lowering the temperature of the shell/fiber interface . The thermal gradient imposed across the interior surface of the liquid layer by this procedure causes evaporation from the surface of the liquid layer. Since the heat of vaporization for the hydrogen isotopes is large compared with the heat of fusion (a factor of 6.6 for 0 2 ) , forced boiling of about 15% of the liquid causes the remaining liquid to freeze on the inner surface of the glass shell. Fig. 3. Solid-fuel Iayer on inside surface of target. D. L. Muslnsld, T. M. Hendenon, R. J. Simms, T. R. Pattinson, and R. B. Jacobs 52 7.0 r - - - - - - - - - - - . . . . . , 6.0 al 70 ·J~m dia. 30 atm bl 100jlm dia . 30 at m cl SO·)Im dia. 100 atm dl100·)1m dia. 100 atm 5.0 E ~ 4.0 TARGETJ LEVEL ..J w ~ 3.0 a: 1- 2.0 d 1.0 100 0 10 20 30 40 50 60 70 80 90 x 10- 3 sec Fig. 4. Piston retraction of radiation shield. 150 190 EXPOSURE TIME x10·3..., Fig. 5. Solid-fuellayer vs. exposure time. Initial fuel approximately 85% of available fuel. Once frozen, the temperature of the fuellayer is held below the melting point of the fuel by conduction throu~h the copper fiber. As a result, the solid layer of Fig. 3 sublimes instead of melting [ ] when the radiation shield is retracted and the target is exposed to room temperature radiation. A fuel layer that must sublime rather than melt has a substantiallifetime, i.e., time during which solid fuel remains as a continuous layer on the inner wall of the target. The amount of solid fuel remaining on the inner surface of this target at the moment the Iaser is fired is determined by the duration of the exposure to room temperature radiation. This exposure time is determined by the retraction speed of the heat shield (Fig. 4). This exposure time also sets some bounds on the range of target size and fuel fill that can be used with this system. Figure 5 shows the amount of fuel remaining as a continuous layer on the inner surface of a target as a function of exposure time, for various size targets and fuel fills. The approximately 30-ms exposure time of the retraction device shown in Fig. 4 allows a targetassmall as 70 J.l.ffi in diameter with a fill pressure of 30 atm tobe irradiated with an appreciable amount of fuel remaining as a continuous layer on the inner surface of the target. Although this technique allows the formation and irradiation of solid-fuel-layer targets, these layers are inherently nonuniform. The liquid (about 15% of the fuel) that vaporizes during formation of the solid layer freezes as a small spheroidal volume localized over the shell/fiber interface. Also, the amount of fuel that sublimes during the exposure time is added to this localized spheroidal mass. Because of this inherent asymmetry these targets are not considered optimal for Iaser experiments. However, integration of the more complex system into the experimental target chamber afforded the opportunity to clearly identify and specify the problern areas involved with timing an active system (the retracting shield) to the Iaser firing sequence. The key requirements or problern areas were then addressed when designing the system for uniform solid fuellayer targets. PRODUCDON OF UNIFORM SOLID-FUEL-LAYER TARGETS There are two separate aspects of producing a uniform solid fuellayer target for Iaser irradiation. First, the system and procedure must provide a source of refrigeration to form the fuellayer in a manner that maximizes uniformity. Second, the system must be amenable to integration of the apparatus and procedures with the experimental target chamber. The constraints imposed by existing systems dictate, to a great extent, the options available. Engineering Aspects of Cryogenic Laser-Fusion Targets 35 53 HFUEL • 2.9a10 6 J HGLASS' 4 .9110 8 J "30 a"' t3 25 0 PELLET DIA !00 ~ 20 300K FILL PRESSUREIPl • IOOATM ..; )• 100~11 ~ ...~ LIQUID+ VAPOR SOLID +LIQUID SOLID Fig. 6. Lifetime of a fuel layer exposed to room temperature radiation. 100 150 200 T I ME , MSEC Achieving uniformity in a solid fuel layer requires that two conditions [8 ] be satisfied. First, any temperature gradient on a given radial surface within the target shell or fuellayer must be small during the condensation and freezing of the fuel. Any temperature gradient within the target is relieved by a redistribution of fuel, evaporation from the warm region and subsequent condensation at the cold region. This redistribution of fuel mass per unit time becomes negligible only below 10 K. Therefore, while the fuel is condensing and freezing (35 to 18 K for the nominal 100 atm, 100-JLm-diameter target), temperature gradients are intolerable. The second condition for achieving uniformity requires that the time for condensation and freezing be minimized. Under the inftuence of gravity the liquid and/or liquid and solid mixture can sag within the target. Minimizing the time minimizes the sag. A condensing and freezing time of 20 ms permits a theoretical nonuniformity of less than 5%. Cooling with gaseous helium can in principle satisfy both of the conditions. The most practical method of minimizing the condensation and freezing time is to heat the fuel in a previously frozen target, vaporize it, and allow it to refreeze rapidly within a cold environment. In a series of well-controlled bench experiments using an isothermal static helium environment, Miller [9 ] has demonstrated the degree of uniformity that is possible with this method. He was able to produce targets that met the criterion of 20% WNU. * In addition, he showed that a CW Iaser is an efficient means of coupling heat to the target. Subsequently, Woerner [ 10] confirmed this result in a similar system. However, integration of a static isothermal environment with the experimental target chamber is a formidable task requiring several Ievels of complexity [ 11 ]. For the existing experimental chamber a static isothermal environment is not required. The primary constraints imposed by the existing experimental chamber are that the main illuminating beams must not be blocked or degraded nor may any of the diagnostic equipment be precluded from Operation. These constraints make it unavoidable that the target be exposed to room temperature radiation for at least a short time prior to the shot. The degradation of the uniformity of the fuellayer when exposed to room temperature radiation is limited for any particular target by both the amount and the physical properties of the fuel. Figure 6 illustrates the point. It shows the calculated lifetime of the various phases of deuterium in a typical isolated target, initially at 10 K, that is exposed to 300-K radiation. (The target is assumed to have an absorptivity of unity for room temperature radiation.) After 10 ms, the * WNU = (maximum thickness minus minimum thickness)/(minimum thickness). 54 D. L. Musinskl, T. M. Henderson, R. J. Simms, T. R. Pattinson, and R. 8 . Jacobs temperature of the fuel has reached the melting point where the uniformity of the fuellayer degrades as melting commences. Thus, forauniform solid-fuel-layer target which is thermally isolated, the maximum exposure time is limited to less than 10 ms. To minimize the amount of vaporized fuel an exposure time significantly less than 10 ms is also desirable. EQUIPMENT REQUIRED FOR UNIFORM SOLID FUEL LA VERS The system that has been designed and constructed to meet the requirements of uniformity and compatibility with the target chamber uses a gaseous-helium shroud and is shown in schematic form, integrated with the target chamber, in Fig. 7. Much like the heat shield, it is lowered from the top of the experimental chamber to envelop the prealigned target and form the fuellayer. Just prior to the target shot the shroud is retracted, uneavering the frozen target. The details of the gaseous-helium shroud are shown schematically in Fig. 8. Simplicity of design is achieved by using a commercially available cryostat and allowing the shroud to continually supply gaseous helium at a constant rate through a pinhole orifice to the region around the target. Four windows are provided at the target Ievel, two for observation and two for a heating Iaser (the CW alignment Iaser of the target chamber) used to vaporize the fuel in situ prior to the rapid refreeze [9 ]. Alignment of the target at the focus of the Iaser system is critical (±5-~tm permissible displacement). One must either attempt to prevent any displacement of a prealigned target while the fuel layer is being formed or provide for repositioning afterward. To avoid the complexity of alignment while the target is under the shroud, the system was designed to minimize target displacement during fuel-layer formation. The target is mounted on a copper post which fits into a collet joint in the extension of the lower cryostat. Before target alignment and formation of the fuel layer, the temperature of the lower cryostat is set at a value less than 10 K to avoid subsequent thermal displacement of the target under the shroud. Although the target is mounted on a copper post as in previous cryogenic experiments (Fig. 1) with CRYOGENIC SHROUD LASER BEAM CRYOSTAT EXTENSION Fig. 7. Gaseous-helium shroud integrated in target chamber. 55 Engineering Aspects of Cryogenic Laser-Fusion Targets COMMERICAL CRYOSTAT GAS FLOW FILTER TARGET LOWER CRYOSTAT EXTENSION Fig. 8. Schematic of gaseous-helium shroud. liquid-layer targets, the target itself is thermally isolated from the post by a 200- to 400-ILm length of 10-ILm-diameter glass fiber. This mounting arrangement prevents target motion due to cooling under the shroud, yet assures that the cold gaseous helium is the dominant source of refrigeration for layer formation. Nonuniformities due to asymmetrical cooling by the fiber are thus avoided. To prevent mechanical displacement of the target, the shroud does not touch the lower cryostat extension. The annular gap between the two acts as a variable impedance to the ftow of gaseous helium from the target region inside the shroud to the rest of the target chamber. For maximum fuel-layer uniformity, the pressure near the target needs to be maximized in order to minimize the condensation and freezing time. This pressure is determined by both the throughput of the pinhole and the size of the annular gap. The throughput into the target chamber is set by the size of the pinhole, which must be chosen so that the pressure throughout the target chamber is no higher than approximately 3 mtorr. At higher pressures, the conductivity of the gas in the target chamber is sufficient to overload the cryostats and prevent operation below 10 K. For this throughput, the pressure near the target is maximized by decreasing the gap to a minimum. Condensing and freezing of typical targets in this system occurs in :::;30 ms. The physical dimensions of the shroud plus the overlap of shroud and targetpost are dictated by considerations of shroud retraction. Exposure of the target to room temperature radiation may be considered to start when the bottom of the shroud reaches the level of the target. From this position the shroud must be lifted clear of the path of the converging laser beams within 10 ms. The acceleration needed to accomplish this can be minimized by maximizing the shroud overlap, i.e., the distance over which the shroud accelerates before exposing the target. The actual overlap chosen (4 cm) is a compromise between maximizing the overlap and assuring that the clearance (1.6 mm) between the inside of the shroud and the target-post 56 D. L. MuiDsld, T. M. Henderson, R. J. Simms, T. R. Pattinson, and R. B. Jacobs CEILING TARGET CHAMBER FLANGE Fig. 9. Schematic of shroud retraction device. extension can be maintained during both initial alignment of the system and shroud retraction. The retraction system is shown in schematic form in Fig. 9. A primary design consideration (recognized when using the piston retraction system of the radiation shield) is the minimization of target motion due to vibration. Since any impulse applied to the target chamber could displace the target, the retraction system is mounted on the ceiling and not on the target chamber. The only mechanical coupling between the retraction device to the target chamber is through a 2-in.-diameter bellows, the minimum size that will permit x, y alignment of the shroud through the vacuum seal (see Fig. 7). The force for retraction is provided by two evacuated 5-in.-diameter bellows. This force is initially counterbalanced by flat-faced electromagnets. Retraction is triggered during the laser-firing sequence by dumping the current in the magneticfield coils. Since the holding force of the magnets is very sensitive to gap, free acceleration of the shroud is achieved quickly and the release time is highly reproducible. Figure 10 shows the position of the shroud as a function of time during retraction. Time zero on the figure is the time when the magnet dumping circuit is triggered. The data shown were taken from multiple exposures on high-speed Polaroid film of an indicator flag attached to the system. The two data points at each time point are the worst-case data for a series of 25 runs conducted over a two-day interval. Total travel is 10.8 cm and occurs in 60 ms. Laserirradiation of the target is timed to occur at the point indicated on the figure before deceleration (provided by two commercial shock absorbers) begins. This timing precludes target vibration due to deceleration shock. With Iaser irradiation occurring at this point, the exposure of the target to room temperature radiation is kept to 4.0 ± 0.35 ms. EVALUATION OF SOLID-FUEL LAYERS To evaluate the quality of the fuel layers produced with this apparatus the interferometer 2 ], shown schematically in Fig. 11, was constructed and used in the e 57 Engineering Aspects of Cryogenic Laser-Fusion Targets ,, 10 9 8 E 7 _,"- 6 > 5 w <( ... er: 4 ~TARGET 3 2 LEVEL 1 0 ~t = :!::350,._ 0 10 20 30 40 50 60 70 TIME x10-3tee Fig. 10. Retraction of gaseous-helium shroud. off-line simulation chamber as weil as in the target chamber. The major feature of this interferometer is the shearing cube C3 ]. With it both high contrast and stability are assured as weil as the ability to continuously vary both the background phase and phase gradient. The illuminating lens is adjusted to focus the Iaser behind the target, producing a diverging beam a few target diameters wide at the plane of the target. The imaging lens is set to focus the crossover of the Iaser beam on the back side of the shearing cube while focusing an image of the target on the image plane. By limiting the focus point on the shearing cube, extraneous phase shifts due to variations in the surface of the cube are eliminated. Phase shift and shearing are controlled by sliding and rotating one-half of the shearing cube relative to the other. so•J. REFLECTIV ITY ARGONION LASER SHEAR ING CUBE Fig. 11. Sheari ng cube interferometer system. 58 D. L. Musinski, T. M. Henderson, R. J. Simms, T. R. Pattinson, and R. B. Jacobs The quality of targets that this system is capable of producing is shown in Fig. 12. Figure 12a is the interferometric image of a DT-filled target when the fuel is in the gas phase. Figure 12b is an interferometric image of the same target when the fuel condensed in the solid phase. Using a criterion based on computermodefing for these interferometric images [9 ] the target shown in Fig. 12b meets the criterion of around 20% WNU. At the present time the ability repeatedly to obtain acceptable uniformity in a solid layer is not entirely satisfactory. A remelt of the target shown in Fig. 12b may produce the target of Fig. 12c as often as it reproduces the image of Fig. 12b. High-speed movie films (500 fps) taken of the vaporize/refreeze cycle clearly show that during condensation, while the fuel is in the liquid state, the fuellayer is very uniform. The nonuniformities observed in the final stable solid layer occur either Fig. 12. (a) Interferometric image of a DT-filled target, fuel in gas phase; (b) interferometric image of a DT-filled target, fuel in uniform solid layer; and (c) interferometric image of a DT-filled target, fuel in a nonuniform solid layer. Engineering Aspects of Cryogenic Laser-Fusion Targets 59 during or after freezing of the liquid. At present not all the factors responsible for the uniformity or nonuniformities are understood. Further study is required to identify and understand these factors so that uniformity of the fuellayer can be controlled and repeated. LASER IRRADIATION OF UNIFORM SOLID-FUEL-LAYER TARGETS The additional operational procedures required for these solid-layer targets have been integrated into the normal routine of the experimental program. The target is installed in the experimental chamber and aligned with the main Iaser, following the same sequence as that used in the liquid-layer experiments [6 ]. First, the gaseous-helium shroud is lowered over the target to freeze the fuel. Next, the CW alignment Iaser, which is coaxial with the main Iaser system, is used to vaporize the fuel within the target. Because of the large angular coverage of the target by the ellipsoidal-mirror illuminating system (see Fig. 12), beam reducers are inserted into t}!e path of the alignment Iaser. These allow the ellipsoidal-mirror illuminating system to focus all the power of the alignment Iaser onto the target through two of the four windows in the shroud. The power of the alignment Iaser is increased until the fuel within the target evaporates. An operator watehing the interference pattern of the target regulates the freezing of the fuel by shuttering the CW alignment Iaser until the uniformity of the solid-fuellayer appears satisfactory. The interference pattern is then photographed and the beam reducers are retracted from the path of the main beam. The final preparatory step is to charge the shroud-retraction device by energizing the holding magnets and evacuating the 5-in.-diameter bellows. Triggering of the retraction device is controlled automatically by appropriate timing circuits locked into the normal automated laser-firing sequence. MIRROR- LENS ILLUMINATION SYSTEM REDUCED BEAM NORMALBEAM BEAM REDUCER '-'· REDUCED BEAM > < NORMAL BEAM Fig. 13. Laser-heating a shrouded target in the target chamber. 60 D. L MuiDIId, T. M. Headenoa, R. J. Süa.s, T. R. Patdaloa,lllld R. 8. Jacobs SUMMARY Experiments with cryogenic liquid-layer targets, using point-contact conduction cooling, have opened the way to experiments with solid-layer targets. The engineering and integration of successively more complex systems identified the critical design features and procedures needed to achieve uniformity of the fuellayer. The technology has been refined and extended in several iterative steps to eliminate the major causes of nonuniformity and to provide an interferometric viewing system that enables an operator to observe and evaluate the layer formation. The gaseoushelium shroud and retraction system were shown to be capable of producing and presenting to the laser a satisfactorily uniform (:520% WNU) solid-fuel-layer target. Successfullaser irradiation of such targets demonstrates that this system now oflers the opportunity to experimentally study in a systematic way a new class of laser fusion targets-uniform solid-fuel-layer targets. ACKNOWLEDGMENTS 'Ibis work was supported in part by the United States Department of Energy under Contracts EY-76-C-02-2709, ES-17-C-02-4149, ED-78-C-08-1598, and DE-AC08-78DP40030. The successes achieved in this cryogenic work are very much a consequence of outstanding teamwork within the Division of Material Seiences at KMS Fusion. The authors owe particular thanks to D. L. Melmoth whose close attention to details avoided many pitfalls and to E. J. Calabro who, despite a long Iist of restrictions and constraints, succeeded in converting a few general concepts into a practical, remarkable, trouble-free design for the shroud-retraction mechanism. Special thanks must also be extended to R. D. Sigler who suggested the use of the shearing cube interferometer and who guided the authors in its use. J. A. Tarvin contributed both valuable discussions and laboratory expertise during the final testing stage of the gaseous-helium shroud. REFERENCES 1. G. S. Fraley and R. J. Mason, Phys. Rev. Len. 35:520 (1975). 2. R. J. Mason, Nucl. Fusion 15:1031 (1975). 3. R. G. Schneggenburger, W. S. Updegrove, and R. L. Nolen, Jr., Rev. Sei. Instrum. 49(11):1543 (1978). 4. T. M. Henderson, R. B. Jacobs, D. L. Musinski, R. J. Simms, and G. H. Wuttke, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 690. 5. T. M. Henderson and R. R. Johnson, Appl. Phys. Len. 31:18 (1917). 6. T. M. Henderson, D. L. Musinski, R. B. Jacobs, and R. J. Simms, in Chem. Eng. Progr. Symp. Series, to be published. 7. R. L. Berger et aL, in Proc. 7th Intern. Conference on Plasma Physics and Controlled Nuclear Fusion Research, VoL 3, IAEA, Innsbruck, Austria (1979). 8. T. M. Henderson, R. J. Simms, and R. B. Jacobs, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 682. 9. J. R. Miller, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 669. 10. R. L. Woemer and C. D. Hendricks, Technical Digest-Topical Meeting on Inertial Confinement Fusion, 'JbE1, San Diego, California, February 7-9, 1978. 11. J. R. Miller, R. D. Day, E. H. Farnum, W. G. Hansen, H. E. Tucker, and W. A. Teasdale, Technical Digest-Topical Meeting on lnertial Confinement Fusion, 'JbE10, San Diego, California, February 7-9, 1978. 12. J. A. Tarvin, D. L. Musinski, T. R. Pattinson, R. D. Siglerand G. E. Busch, in Proc. 23rd Intern. Symposium and Instrument Display of the Society of Photo-Optical Instrumentation Engineers, San Diego, California, to be published. 13. B. J. Sanders, Appl. Opt. 6:1 (1967). B-1 ENERGY TRANSFER IN A SYSTEM OF SUPERCONDUCTIVE MAGNETS M. Masuda, T. Shintomi, and K. Asaji National Labaratory for High Energy Physics Ibaraki, Japan INTRODUCTION Magnetic fusion reactors, equilibrium coils, and ohrnie heating coils require large pulses of energy which are repeated over relatively short periods of time. Accelerator magnets fabricated with normal conductors or superconductors have similar operating conditions, the same pulse time, and require almost the same amount of energy, namely, hundreds to thousands of MJ. These large pulsed energies can a:ffect apower grid adversely. Many schemes have been proposed for accelerator magnets; some have been constructed and operated successfully. One scheme entails reactive power control by thyristors C] for a proton synchrotron; however, no matter how precisely the reactive power has been controlled by the thyristor device, the active power has never been properly controlled. In fusion reactors, energy storage devices are needed not only to suppress adverse effect to the power grid but also to improve efficiency by saving energy losses over an operation cycle. Schemes also have been proposed for fusion reactors, such as homopolar generators, flywheel motor generators, and superconductive energy storage. The flywheel motor generator was employed in older accelerators; but for the past decade, it has not been employed in modern accelerators due to mechanical failure of the generator pole. Now the static rather than the rotating machine has become conventional. Minimizing the amount of energy from the power grid is a problern common to both fusion reactors and large accelerators. Pulsed superconductive energy storage is one of the solutions to this problem. Numerous energy transfer methods have been proposed; the most promising The method proposed by the Wisconsin group ones have already been reviewed is one of the most useful methods to transfer a large amount of energy between coils. The purpose of this paper is to describe a technique useful for superconductive energy storage, namely, the energy transfer between two superconducting coils, assuming that one is the storage coil and the other the accelerator magnet or the poloidal coil in a tokamak, for example. eJ eJ. 61 62 M. Masuda, T. Shintomi, and K. Asaji Fig. 1. Placement of two 100-kJ superconducting coils. EXPERIMENTAL PROCEDURE Two superconducting coils, each storing 100 kJ, were placed in the same Dewar. A similar experiment has been already carried out by Argonne group [4 ] where one shot energy transfer has been attempted. In the present experiment it has been demonstrated that not only is it possible to transfer pulsed energy from one coil to the other but also to allow a surge of energy to flow back and forth between the two superconducting coils. Figure 1 shows the assembly of the two 100-kJ superconducting coils. The specifi.cations for the coil are given in Table I. The circuit and the specifi.cations for the experiment are shown in Fig. 2 and Table II, respectively. Table I. Specifications of the 100-kJ Coil Dimension of coil Inner diameter, m Outer diameter, m Length, m Number of turns lnductance, H Dimension of wire, mm 2 Diameter of filament, p.m Number of filaments Composition of filament Cu ratio Twisted pitch, mm Critical current, A 0.2 0.25 0.2 1284 0.22 2.4 X 1.53 55.6 271 Nb-Ti 4.16 29.1 900 at 5 T 63 Energy Transfer in a System of Superconductive Magnets CO I L 1 BRIDGE 1 COMM UTATION CAPACITORS BRIDGE 2 COIL 2 Fig. 2. Schematic diagram of test arrangement. Control of the energy transfer rate was accomplished by controlling the phase delay from one bridge to the other. The circuit was completely symmetrical, and the reverse of the phase delay inverted the transfer. The phase angle control circuit is shown in Fig. 3. The high frequency signal (up to 1 MHz) from the voltage-tofrequency converter (VFC) was fed to two dividers and to the data distribution circuits for the thyristor bridges. The phase diflerence between the two bridgeswas controlled by input fed to a digital comparator by means of an analog-to-digital converter (ADC). The sign for the input data depended upon whether there was a phase advance or delay from one bridge to the other. The transferrate and direction could be controlled by the input signal. The signal from the VFC was changed by input which was proportional to the current ftowing in the commutation capacitors. The output from the clock determines the frequency of the ac voltage between the capacitors and then the control voltage of the capacitor terminals for all conditions. Figure 4 shows the waveforms of the ac voltage and dc output voltage for various phase angles of the bridges. The experimental procedure for each runwas as follows. A dc power source (50 A and 100 V) was used to drive thyristor bridge 1 and this generated an ac voltage of 400-500 Hz between the terminals of the commutating capacitors. The generated ac voltage was rectified and then coil 2 was charged up to its maximum by thyristor bridge 2. After this charging cycle the system was isolated Table II. Specifications for the Power Supply Input dc power supply Voltage, V Current, A Thyristor for the bridges Reverse voltage, V Rated on-state current, A Commutation capacitors Capacitance/arm, pF x 5 Rated voltage, V Connection 100 50 1200 1000 500 200 star 64 M. Muuda, T. Shintomi, and K. Asaji DATA INPUT f DATA INPUT 9 Fig. 3. Control circuit for the phase angle of the bridges. from the dc power source by a switch and the stored ener gy in coil 2 was transferred immediately to coil1 by the inverse Operation of bridges 1 and 2. Afterall the energy had been transferred to coil 1 by the thyristor bridges, the process was reversed and the energy was transferred back to coil 2 by the inverse operation of the thyristor bridges. The main feature of this experiment was the possibility of obtaining a continuous back and forth ftow of energy between the two coils. However, the dissipation of energy particularly in the thyristors limited the process to about five to seven cycles. Figures 5a and 5b show the recorder traces for the one-shot energy transfer and the continuous oscillation of energy between the two coils, respectively. ESTIMATION OF LOSSES Estimation of losses, one of the most important factors in evaluating energy transfer, has been attempted for the one-shot scheme. Schematic diagrams of the circuit and the current waveforms of the two coils for this calculation are shown in (A) (8) Fig. 4. Waveforrns of ac voltages and dc voltages for various phase angles. Legend: (A) 0°; and (B) 90°. 65 Energy Transfer in a System of Superconductive Magnets 500 (A) 400 <{ 1- z 300 IIJ """" ::> u ...J 200 8 100 00 10 5 15 TIME , SEC 500 400 <{ 1- zIIJ """" ::> u ...J 300 200 8 100 0 0 15 10 5 TIME , SEC Fig. 5. Current waveforms for energy transfer. Legend: (A) one-shot energy transfer; and (B) energy transfer using two coils. Figs. 6a and 6b, respectively. The energy in this analysis is transferred from IP to I. in a period of time T. Since the two supercondueting coils were identical, their nominal rated currents were the same. In practice the actual current was about 10-30% smaller than the transmitted current because of losses. These losses were mainly caused by the thyristor forward voltage drops, the protection resistors and the circuit cables. The energy loss per one-shot energy transfer can be estimated as follows: Ä r[ W= nip-;Ä V+ r( Ip-;) + ni.-;Ä V+ r( I.-;) dt =~LI; -~LI; The current ratio, I./ IP = 2 2 ] (1a) (1b) a, is obtained by solving equation (1). In the present case, M. Masuda, T. Shintomi, and K. Asaji 66 ~T~ Fig. 6. Schematic diagram of circuit and current waveforms of the two coils. L 1.0 lp : 300 A ~ ~ 0. 5 T' 1. 0 SEC IL p -- 500 A 0.5 'T , SEC L lp : 700 A Fig. 7. Calculated efficiencies for the one-shot energy transfer with three different amounts of current. Legend: - - , without resistors; -- -, with resistors. ? 0.5 T, SEC Energy TrlliiSfer in a System of Supercondnctive Magmets 67 L, r, and a V were assumed tobe 0.22 H, 10 mfi, and 2.0 V, respectively. Substitution of these values into the above relations yields ( 0.11 + -10- T) 2 3 a 2 2 1 ( 0.11- 2T-1 10+ 3T-a- T) = 0 Ip Ip 3 (2) Solving for a results in T (3) a = 1-231- 0.03T p On the other band, IP and I. are shared by the coils and the protection resistors as shown in Fig. 4. The sharing ratio, ß, is given as L ß = Ip/ I P = 1 1 + To/ T ( 4) r Thus, the total efficiency of energy transfer, Tl = (I~'/ I!' )2 , is obtained from "' = [ß 2 ( 1 _ 23T I;ß _ 0.03T) (5) The calculations have been performed for three different coil currents and operating with or without the protection resistors. The results are summarized in Fig. 7. EXPERIMENTAL RESULTS AND DISCUSSIONS More than 30 experimental runs have been carried out. Transfer efficiencies, "'' have been calculated and good agreement with the calculated results has been obtained as shown in Fig. 8. Since the coupling between the two coils is less than 5%, the efficiencies were only slightly affected. The shorter the transfer time, the better " WITHOUT RESISTCRS o WITH RESISTCRS CALCULATION t; i5 u u: 0.5 lL UJ o~----~-----L----~L-----~-----L- 0 2 3 4 5 TRANSFER TIME , SEC Fig. 8. Experimental results for the efficiencies of one-shot energy transfer with and without protection resistors. 68 M. Muada, T. Sldatoaü. lllld K. Auji the transfer efficiency becomes when the protection resistors are removed. When the protection resistors werein the circuit, quick transfer of energy became difficult. One of the main reasons for a low transfer efficiency could be the protection resistors. Even if varistorsbad been employed for protective purpose, they might also have bad an adverse eflect on the quick transfer of energy. The decay of the total energy in the system is caused by the Iosses of the thyristor, Iead wires, and the commutation capacitors except for the protection resistor loss. The fact that the thyristor loss is dominant is explained by the diflerence of the decay time constants in Fig. 5 as follows: only one bridge is related to the decay in Fig. Sa. On the other band, two bridges are related to the faster decay in Fig. Sb. The diflerence of the decay was also confirmed by an estimation of the losses. The thyristor loss results from the forward voltage drop which is a constant value for any type and size of thyristor. Therefore, the thyristor loss is linearly proportional to the current. On the other band, the stored energy is always proportional to the square of the current. Thus, the thyristor loss for a larger system, such as 100 MJ or more for example, will play a minor role in the Iosses of a superconductive energy storage system. NOTATION Ip = primary current flowing into the commutation capacitors I} = primary coil current I, = secondary current flowing into the commutation capacitors L = inductance of the coil n = number of thyristors R = resistance of the cable r = resistance of the resistor t = time during the energy transfer T0 =time constant between the superconducting coil and the protection resistor, L/ R 4 V = forward voltage drop 4 W = energy loss per one-shot energy transfer a = current ratio, Ij IP ß = sharing ratio of the currents, IJ It; TJ = transfer efficiency T = transfer time REFERENCES 1. M. Masuda, S. Matsumoto, and T. Shintomi, IEEE Trans. Nucl. Sei. NS-24:1306 (1977). 2. R. S. Ramshaw and E. P. Dick, in Advances in Cryogenic Engineering, Vol. 21, Plenum Press, New York (1976), p. 149. 3. H. A. Peterson, N. Mohan, W. C. Young, and R. W. Boom, in Proc. Intern. Conference on Energy Storage, Compression and Switching, Plenum Press, New York (1974), p. 309. 4. R. E. Fuja, R. L. Kustom, and R. P. Smith, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 97. B-2 THERMAL CYCLE TESTS OF A MODELED SUPERCONDUCTING TRANSMISSION LINE C. F. Sindt and P. R. Ludtke NBS Thermophysical Properfies Division Boulder, Colorado INTRODUCTION Superconducting power transmission lines (SPTL) will be enclosed in an insulated conduit which typically is a vacuum insulated, double-walled vessel. The outer wall is rigid and the inner cold wall has provisions for thermal contraction and expansion. Three cable assernblies (one for each phase) are used in the alternating current systems. In the system designed by Brookhaven National Laboratories (BNL), high-pressure cold helium gas at 6 K ftows down the hollow core of each cable. This gas is recooled to 6 K from 8 K by expansion at the cable end and ftows back around the outside of the cables. In the Brookhaven installation, the ends of the SPTL cable will be fixed (i.e., the length of the cable held constant) so that a significant thermally induced axial stress will develop in the cable as it is cooled. The goal of the testswas to investigate the thermoelastic behavior of a 3.5-m section of a model of a SPTL cable constrained to constant length while the cable section was subjected to repeated thermal cycles between ambient and operating temperature. The cable, which was constructed according to BNL specification number PTP # 76-22, was built up from a number of helically wrapped layers (bronze core, superconductor, dielectric); one of the outer layers is a Iead sheath which provides a hermetic seal (see Fig. 1). The cooldown tensile Ioad produces an elastic deformation in the helically wound elements which act as coil springs. If friction between these elements is neglected, this spring Ioad is small compared to the Ioad on the Iead sheath which is stressed beyond its elastic Iimit during cooldown. Likewise, the Iead sheath will be compressed beyond the elastic Iimit during warmup. Although repeated thermal cycling of an operating SPTL is not expected, it is nevertheless necessary to provide for this contingency. Thm~ the cable must remain functional throughout a number of thermal cycles. The primary test objective was to observe if physical darnage to the cable would occur as a result of thermal cycling. Additional objectives were: to determine the cable tensile and compressive Ioads during thermal cycling; to test the ability of the end clamps to withstand thermal and assembly cycles; and to determine if the cable or end clamps developed leaks at operating temperature and pressure. To accomplish these objectives, two different 3.5-m lengths of cable were subjected to a number of thermal cycles while maintaining the cable length constant. 69 Bronze Co re, 2 He 1 i c e s 5 Strips Alumi num Sta bilizer, 2 Hel ic e s , 5 strips Coppe r Supe r cond uctor, 2 Helic es 5 St rips Inne r Die 1ec t ri c Screen ( in tercalate d) , 2 Tapes Me t al l i zed Po l y i mide Oi e l e c tr ic , Polye thyle ne 93 Laye r s Oute r Oie lect r ic Scree n (i nt e rca l a t ed), 2 Tapes Coppe r Superco nductor, 2 Hel ic es Al umi nu m St abilizer, 2 Helic e s Outer Conductor Jnsulatio n ( kap t on), 4 layers St ai n less Steel Comp r ession layer Taped Fiber9lass Thermal lnsu l ation, 2 layers Lead Gas Jacket Armer, Myl ar- Brgnze la~...ent l ay Angle of 80 to 85 Dimensions in Millimeters Fig. 1. Cut back view of cable. 0 D l L Di rect ion Thi ckness A 21. II 120 RH 1.02 8 23.11 133 LH 1.02 c 23.88 89 RH . 38 D 24 . 64 36 LH .38 24.84 112 RH . 10 F 25 . 04 112 LH . 10 G 25.15 H 40.38 RH . 20 7.62 I 40.48 LH .05 J 40.69 185 RH . 10 K 40 . 89 185 LH .10 L 41.65 ! 57 RH . 38 !57 LH . . 38 RH .51 M 42.41 N 43 . 43 0 44.45 p 45.46 . 51 Q 49.79 2 .16 R 50 . 14 . 178 50.50 . 178 s 229 RH .51 Thermal Cycle Tests of a Modeted Superconducting Transmission Line 71 EXPERIMENTAL APPARATUS The test fixture, Fig. 2, must meet several requirements in these constant-length thermal cycling tests. The test requires (1) a mechanism for maintaining the cable length constant with varying Ioad; (2) aDewar and cryogenic fluid system for cooling the cable to liquid nitrogen temperature; and (3) a gripper and seal arrangement to transmit Ioad from the cable to the test fixture while maintaining a hermetic seal with the lead sheath. The test fixture, which accommodates a 3.5-m cable segment, has a reetangular steel framework built up from two 100-mm-square tubes, 5.4-m long, that are bridged at each end by 75 x 127 mm H beams, 0.5 m long. A screw mechanism at each end may be manually adjusted so as to maintain the cable length constant in the face of variable loads. The cable is enclosed in a polyurethane-foam-insulated spool piece and the end clamps are enclosed in the vacuum-insulated housings attached to each end of the spool piece. The cable, spool piece, and end housings are all supported at either end by rails with linear ball bearings (see Fig. 2). The cable assembly is isolated from forces in the support structure by means of internal and external bellows in the end housings, as weil as by means of 0-ring slip joints at the end housings. The lead seal and terminal clamp assembly for attaching the end of the cable to the test fixture and the assembly procedure were provided by Brookhaven National Laboratory. An enclosure cap was added to the clamp assembly, which is shown in Fig. 3, to facilitate pressurization of the inner line and to provide the structurallink between the cable and the test fixture. INSTRUMENTATION The length of the cable is monitored with 30 power microscopes which are aligned with view ports 86 mm from each end of the spool piece. These microscopes are supported from the concrete ftoor by heavy duty tripods, so their position is independent of the Ioad on the cable and test fixture. The cable length reference points are scribe marks on Invar rods which are attached to the end fixtures. The cable length was held constant to within ±0.12 mm using this system of microscopes and adjusting screws. Cable temperatures were determined with type-T thermocouples referenced to a boiling liquid nitrogen bath. The cable was cooled from one end rather than uniformly along its length; the purpose of the temperature measurements was to determine when the entire cable segmentbad cooled to the desired temperature. The measurements were not intended to represent the temperature of the cable composite since the sensor could not be placed in the body of the cable without destroying the integrity of the Iead sheath. The axial tensile or compressive Ioad on the cable was determined with a Ioad cell [2000-lb (9000-N) capacity]. A purely axialload at the cell is assured by the guide mechanism for the cable extension tube shown in Fig. 2. The calibration accuracy of the Ioad cell is 0.25% of full scale. The precision of the load cell was checked with dass S weights; the measurement uncertainty was 0.1%. The axialload on the cable differs from the force on the Ioad cell and is given by the expression (1) where Fe is the force on the cable, Fd is the force measured by the Ioad cell, FP is the 0 Fig. 2. Test apparatus. Legend: A, Liquid fill and pressure port for i.d. of cable; B, cable length adjusting nut; C, subassembly end guide and supports; D, liquid fill for o.d. of cable; E, auxiliary supports (used only when necessary to remove vacuum dewars on ends); F, observation port; G, base axis of symmetry; H, vent for o.d. of cable; I, Observation port; J, subassembly end guides and supports; K, Ioad cell; L, vent for i.d. of cable; M, cable length adjusting nut. 0 0 - -- @ I 1 ·---·---·---·---·--·--·--·-- --:-·--·-- 5384.8 ± 3 mm (212 t 1/8 in. = 17.66 ft) - - - - - - - - - - - - -------, r--- - - - - - 3048 .0 ± 1 mm (120 ± 5/ 128 in .= 10ft)- - - - - -- Vscuum Invar Spscer Fig. 3. E nd clamp. Indium Ses/ Sleeve ~ ~ i I f IICI I I • ia. 0.. !) i .. ~ 74 C. F. Slndt and P. R. Ludtke Fig. 4 . End assembly schematic. pressure force due to the differential pressure across the end fixtures, Fcp is the axial force on the cable resulting from internal pressurization, and F1 is the friction force on the cable and cable extensions. The friction force from the supports and 0-ring seal is estimated to be less than ±45 N (± 10 lb )-small compared to the other forces . The pressure force, Fm is due to the differential pressure across the inner bellows assembly. A tensile force of 151 N (34 lb) results. In considering the effect of internal pressurization of the cable, it has been observed that the behavior of a composite cable can be much less predictable and reproducible than the behavior of a noncomposite cable. Referring to Fig. 4, the axial force on the composite cable due to internal pressurization is Fcp = P[A.- (A. - Ac)(l- a)C] (2) where Pis the differential internal pressure, A. is the inside area of the Iead sheath, Ac is the inside area of the bronze core, a is the void fraction of cross-sectional area of the composite inside the Iead sheath, and C is the degree of coupling between the inner composite and the Iead sheath and clamp assembly. For an internal cable differential pressureof 1.5 MPa, Fcp mayvaryfrom 2470 N (555lb) to 340 N (77lb), depending on the voids in the cross-sectional area of the inner composite and the degree of coupling between the Iead sheath and the inner composite. Another important consideration is the effect internal pressurization has on the spring constant of the various metal helices in the cable. The bronze core and other metal helices may exert axial forces as a result of temperature change, and these forces may change as the result of internal pressurization. This is because the various helices behave as springs in the absence of friction , but could tend to behave as rigid tubes if there is friction between them. Such friction could decrease if the Iead sheath were pushed radially away by pressure, or increase if the sheath contracted more than the rest of the cable during cooldown. TEST PROCEDURES Two cable samples were tested, each by a different procedure. Cable No. 1 was cooled and then pressurized, whereas cable No. 2 was cooled with the core pressurized. The latter procedure simulates the expected operating conditions more closely. Cable No. 1 was first leak checked by pressurizing the cable interior with helium gas to 1.5 MPa and measuring the rate of pressure decay in the valved-off system over a 15-min period. The sensitivity of the pressure measurement was 3.4 kPa. All Thermal Cycle Tests of a Modeled Superconducting Transmission Line 75 pressurization and depressurization cycles were made in 345-kPa increments, so that the cable could be maintained at a fixed length by means of the adjusting screws. After the leak check, the pressure was bled down to ambient pressure and the cable was cooled to liquid nitrogen temperature by ftowing liquid nitrogen through the core and around the outside of the cable. The liquid nitrogen ftow rate was adjusted so as to achieve cooldown in about 90 min. The cable was then again leak checked by measurement of the pressure decay using helium gas prechilled to about 80 K. Following the leak check, the cable interior was bled down to ambient pressure (again in 345-kPa increments to allow for cable adjustment) and the cable was warmed by ftowing 350 K nitrogen gas through and around it. The cable warmed to 200 K in about 3 hr. The remainder of the warmup was achieved unattended in the following 12 hr using a low ftow rate of ambient temperature nitrogen gas. The procedure for cable No. 2 was the same except that it was cooled down by ftowing liquid nitrogen around the outside of the cable while the interior was pressurized to 1.5 MPa with helium gas. Cable No. 1 was cycled twenty times and then removed for inspection. Cable No. 2 was cycled five times, removed for inspection, then reassembled and cycled an additional fifteen times. Both cables were preloaded with 590-N (134-Ib) tension [440-N (100-lb) Ioad cell and 150-N (34-lb) bellows force] before the firstthermal cycle. The cable length was then referenced by fixing the microscope cross hairs on the scribe marks of the Invar rods, and this length was maintained throughout the test series. RESULTS AND DISCUSSION Cable No.l Figure 5 presents Ioad cell vs. time for four typical thermal cycles of the 20 cycles to which cable No. 1 was subjected. The test data are constrained to definite time intervals, i.e., actual cooldown time may have varied slightly (± 10 min) from the 1! hr indicated. Following the first thermal cycle, the preload on the cable changed from the initiai440-N (100-lb) tension (Ioad cell reading) to 1330-N (299-lb) compression and remained near this value for the remaining 19 tests. This shift in the initial cable (Ioad cell) Ioad occurred with both cables and is attributed to the yielding of the Iead sheath which occurs during the first thermal cycle. As the cable is cooled the Ioad passes through a maximum and then decreases slightly as the cable comes to equilibrium at 7 6 K. This behavior may be due to creeping of the spirally wo und core with respect to the Iead sheath. As the cable is pressurized, the Ioad cell force is reduced by approximately 4000 N (900 lb) compared to the maximum of 2470 N (555lb) that is calculated for the pressure times area term. Fora solid cable the two forces must be equal. In this case, however, it is likely that pressurizing the Iead sheath releases the grip of the sheath on the inner composite sufficiently that the composite relaxes even though its length is held constant. A recovery of 800 N (180 lb) in the tension Ioad occurs as the cable core is depressurized. This is within the range of 428 N (961b) to 2466 N (554lb) which may be attributed to the pressure times area term, but differs appreciably from the average change of 1600 N (360 lb) which occurred when the cable was pressurized for leak checks at ambient temperature. The hysteresis in these ambient temperature pressure cycles was small. The Ioad decreased by 1646 N (370 lb) with pressurization and increased by 1557 N (350 lb) with depressurization. After test number 20, the cable was removed from the test fixture. The enclosure cap and split-end cap were removed from the clamp assembly, revealing "" 0 800 1000 0 ~-800 0 E "'~"' -600 .~ -400 0 - 200 '-- olJ 200 ~ 400 !:. ~ 7 600 w -' -' u 0 "' "- u I- 1400 I- 1600 1800 2000 2200 I- t 2400 I- J t 0 V I V l .b I o I I I I I I I I I I I I I I D l cf> 8lf 20 min PressureLoss Leak Check a t 76 K to ""(1> JD "'00 v.-, 0 a> vo ~ ~ ~~~~ '\JO VdJ Coo l to 76 K 2 I I I I I f 1 CR l "~o D "' V öl A 0 "'o 0 0 l> D 0 01\. ~larmup Period, 0 ~ CXl TIME, hrs oR ov A OV V 4 v o o0 " v Circulate Warm N2 Gas Thru & Around Cab le v Do oo v ~. Fig. 5. Force vs. time for cable No. 1. t.O "dl~ oB ( Oepress uriz e II II I .. Unattended Overni te • ~ ~ I •I I I 0 ~ Cl> " oo ~ D I I - -24 "bo 'SO 0 ij ~0 ~ I I I I I I Oepressuri 1\I~I 1 1 1 I I J/ I ~ Pressure ss Leak Check Ambient Temp . 20 - min Pressur i ze I nside of Cable to 15 atm o - Test M20 o - Test #13 Test Ml2 v - Test #3 l> - -2000 0 2000 4000 6000 8000 10,000 ..J 0 0 u w .. -' ..J 0 u "" "' "- "'c0 ... .,"' z: . Thermal Cycle Tests of a Modeted Superrondoding Transmission Line 77 Fig. 6. End view of cable No. 1. tbe end of tbe cable inside tbe stainless steel strands. This examination sbowed tbat tbe dielectric and outer dielectric screen bad unraveled sligbtly as sbown in Fig. 6. This was probably due to tbe relative movement between tbe cable layers near tbe severed end of tbe cable. After removing tbe cable from tbe center spool piece, it was noted tbat tbe Iead sbeatb bad extruded radially outward between tbe inner bronze wrap and bad crossed tbe outer bronze wrap, as sbown in Fig. 7. This radial bulge or spiral annulism occurred at tbe end opposite to wbere liquid nitrogen coolant was introduced. In spite of tbe spiral annulism, tbe Iead sbeatb remained intact, and tbere was no leakage in tbis area. Fig. 7. Annulism in cable No. 1. 78 C. F. Sindt and P. R. Ludtke Cable No. 2 In view of the problern with cable No. 1, the test procedure was reviewed and it was decided that cable No. 2 should be cooledunder pressure since this is the manner in which the cable will be cooled in practice. lt was also decided to thermal cycle the cable five times and then remove it for inspection before completing the full 20-cycle test. Inspection after five thermal cycles showed that the lead sheath of cable No. 2 bad expanded outward slightly at the "windows," where the Iead sheath was covered only with Mylar. Otherwise the cable bad the same appearance as when it was installed in the test fixture. This same "window" expansionwas observed with cable No.l. The cable was reinstalled in the test fixture and testing continued for another fifteen cycles. Again, no perceptible darnage was observed other than the slight expansion at the "windows," which was noted after the first five cycles, and a waviness in the surface which became noticeable when sighting along the cable. The outside diameter of the cable was measured at random locations along its length in order to determine the magnitude of the surface irregularities. From 18 different measured diameters in the two planes, the maximum difference was 1.55 mm (0.061 in.). Cable No. 2 showed much less evidence of the dielectric and outer dielectric screen unraveling or fraying as shown in Fig. 6. Load cell vs. time measurements for five of the last fifteen cycles is shown in Fig. 8. During a typical test on cable No. 2, pressurization of the inner line (ambient temperature) resulted in a decrease in the load cell force of 2220-2670 N (500600 lb) as shown in Fig. 8. During cooldown the Ioad cell force changed from approximately 2670 N (600 lb) in compression to 7120 N (1600 lb) or 7560 N (1700 lb) in tension. Cable No. 2 was then warmed in the same manner as cable No. 1 (no internal pressure). The resulting Ioad cell force the next morning, with no internal pressure, averaged 670-N (150-lb) compression for the first five cycles and approximately 1330 N (300 lb) for the following fifteen cycles. Both cables were installed with a tensile force of 440 N (100 lb) (no internal pressure) but after the first cooldown cycle, this changed to 1110-N (250-lb) compression for cable No. 1 as shown in Fig. 5 and a Ioad cell force of 2670-3110-N (600-700-lb) compression for cable No. 2 as shown in Fig. 8. The Ioad cell forces on cable No. 1 given in Fig. 5 are considerably higher than those on cable No. 2 in Fig. 8 because the latter was internally pressurized. GENERAL DISCUSSION There was no measurable leakage across the Iead sheath or the Iead seal clamp assembly during any of the 40 tests. The cause of the annulism in the Iead sheath of cable No. 1 is not clear. lt could be the result of buckling due to the compressive forces that occur on warmup. If some yielding due to pressure forces bad already occurred at that "window" site then it would be more susceptible to buckling. Whether or not the change in the operating procedure prevented cable No. 2 from developing an annulism is problematical, since the mechanism of failure for cable No. 1 is unclear. lt could be fortuitous that cable No. 1 buckled and cable No. 2 did not. On the other band, it is possible that cooling the cable under pressure prevented the annulism by either (1) maintaining tight contact between the Iead sheath and the outer Mylar-bronze wrapping, thus preventing relative movement and random "' 0 ""--' 0 u --' _. .., ~ u 0 0.. "' ::r: Vl .., Vl z 0 e c: Q.l .. .&:. ·.- IQ.)U .. ... o..v 1-.>< ~ .."' .....' .o .. c:: I.. .....I IQ,) j ' I I~ ~~ o..- ·- !r..U"'' ~ ~~3~ ·.- .. .. .. ::3 z: N Vl .. 0 z ,_ u "" "" .0 2000 c X 1 Cold Pres sure- \ Loss Leak Checkl Cooldown t o 76K ~" ttJ9 t.~ 9 ~ • .p o. t. o .·~'11 t. o'B zel 01 I' 0 0 • 9 0 0 9 6 9 ~20 c #18 • 9 9t. 0 • "v 9 0 • t.~9 • Ambien t Temp. ~n l owing t-tornln Fo 1 1 pr essure (no interna 0 Una ttended Ove rn i te Warmup fl ~ • I ' ~ I ~ Cable Stress a~ Warmup Period , Ci r cu l a t e Warm N2 Gas Thru And Ar ound Cab 1e T I ME , hrs 0 • 6 t8 0 t. 113 16 • Coo 1down Test Fig. 8. Force vs. time fo r cable No. 2. 0 • • 9 .,;:. t t. :.. Oepress uri Inner Line -2000 6000 .." "" u 0 0 ""--' u "" --' .... 0 "' u... z "'c:0 ... ~ ft g· i- ~­ i i IIQ I I i. • i j I ;I 80 C. F. Siadt aacl P. R. Ltultke Table I. Etled of Cable Pressure on Load CeU Reading Change in Ioad cell force, N(lb) Wann Pressurize to 1.5 MPa Depressurize to 0.1 MPa Cold Pressurize to 1.5 MPa Depressurize to 0.1 MPa Cable No. 1 Cable No. 2 -1650 (-370) +1560 (+350) -2220 to -2570 (-500 to -600) -4000 (-900) +800(+180) +440(+100) enlargement of the window, or (2) allowing freer movement of the composite with respect to the sheath, and thus reducing the compressive Ioad that the composite exerts on the sheath during warmup. The thermoelastic behavior of the cable is quite complex as evidenced by Figs. 5 and 8. In particular, note the change in Ioad cell force as the cable is pressure cycled between 0.1 and 1.5 MPa as summarized in Table I. Without a detailed stress analysis of the complex cable structure, one can do little more than speculate on the cause for such differences in the change in the Ioad cell force. lt would seem clear, however, that the thermal-pressure history, and how this history affects the coupling between the shield and the core, has a large effect on the force transmitted from the cable to the fixtures. B-3 DESIGN OF A 400-kJ PULSED ENERGY STORAGE COIL* S. K. Singh, C. J. Heyne, D. T. Hackworth, M. A. Janocko, P. W. Eckels, and J. H. Murphy Westinghouse Electric Corporation Pittsburgh, Pennsylvania INTRODUCfiON The design of the conductor for the 400-kJ, 25-kA pulsed energy storage coil Cl, being built by Westinghouse for the Los Alamos Scientific Laboratory, differs n significantly from the conductor fabricated previously for a 300-kJ coil because of the low-loss requirements and fast pulse rate. As a consequence of the short discharge interval and low-loss requirements, the conductor in this application has a low copper-to-superconductor ratio. DESIGN CONSIDERATIONS The superconducting energy storage coil specification under discussion required an energy storage of 400 kJ at an operating current of 25 kA with a discharge time of 0.7 ms. In addition, the coil structure is required to withstand the forces resulting from a current of 35 kA, and the loading conditions associated with a ten-coil stacked assembly. The required coil inductance is 1.28 mH resulting in a design terminal voltage of 60 kV for the 0.7-ms discharge. The superconductor current density should exceed two-thirds of the short sample value measured along the Ioad line with cooling provided by immersion in a liquid helium bath. The combined energy Iosses during a full charge-discharge cycle from all sources including eddy currents, self-field, hysteresis, and mechanical motion must be less than 0.3% of the 400-kJ stored energy when the coil is treated as a centrally located coil in a solenoidal stack of ten such coils. Because of the inductance and size constraints and the desire for a fully ventilated coil, it was decided to adopt a multilayer, helically wound coil design. The resulting design utilizes a stranded and cabled conductor in a three-layer configuration, with an open structure and pool boiling. The loss requirement necessitated a nonconducting structure for conductor support. The major coil parameters are shown in Table I. * Work supported by the Los Alamos Scientific Laboratory, University of California, under Contract No. X67-6665. 81 82 S. K. Siap, C. J. Heyne, D. T. Hackworth, M. A. JIIIOdto, P. W. Eckels, and J. H. Murphy Table I. Parameters for 400-kJ Energy Storage Coil Inductance, mH Stored energy at 25 kA, kJ Length, cm Diameter, maximum, cm Number of layers Centtal field at 25 kA, T Matrix Superconductor Current, kA Current, fault mode 0.1 s duration, kA Voltage across coil, kV Operating point of superconductor Discharge period, s Charge period, s Holding time, s Cycle time, s Energy Iosses full cbarge--discharge cycle, % Cooling mode e1 1.2s 400 ± 12 -71 67.47 Odd, 3 preferred 2.2 to 2.5 [4 ] Cu-CuNi NbTi 25 35 60 >~ along Ioad line 0.0007 to 1.0 30 to 300 30 to 300 300 to 900 < 0.3% Pool boiling, 4.5 K CONDUCTOR DESCRIPTION The cross section of the superconducting cable is shown in Fig. 1. It consists of 12 primary cables cabled around an elliptical mandrei made from fiberglass, stainless steel, and Kapton insulation. The winding pitch for these cables is approximately 28 cm. Each primary cable consists of seven subcables which are wound around a fiberglass mandrel. The winding pitch for the primary cables is 2.54 cm. PHOSPHOR BRONZE DUMMY COO DUCTOR I>Ti Cu Fig. 1. Cross section of 400-kJ coil conductor. Design of a 400-kJ Pulsed Energy Storage Coil 83 Table II. Cable Characteristics Superconductor characteristics Peakoperating current density, iscop• A/m 2 Finished strand diameter, 2R 0 , mm Outside radins of superconducting bundle, R 1, mm NbTi filament diameter, d, p.m Diameter of copper sheath, d 2 , p.m (across hexagonal flats) Diameter of cupro-nickel sheath, d 3 , p.m Filaments per strand Filament twist pitch, mm Overall Cu 0 .9 Ni0 . 1 : Cu: NbTi ratios Interior Cu 0 .9 Ni 0 . 1 : Cu: Nb Ti ratios Omega insulation thickness, p.m Number of strands per cable Volume per cable, cm 3 3.175 1: 1.5: 1 0.36: 1.5: 1 25.4 504 6722 Dummy conductor characteristics Finished strand diaml:ter, 2R 0 , mm Composition Number of strands per cable Volume per cable, cm 3 0.381 Phosphor-bronze 84 1120 Mandrei characteristics Width, 2a, cm Thickness, 2b, mm Composition Volume per cable, cm 3 1.54 3.81 304L 6850 1.52 X 109 0.381 0.172 14.4 21.7 23.2 199 The subcable is fabricated from six insulated mixed-matrix superconductor strands cabled around a phosphor-bronze dummy conductor. The winding pitch for the subcables is approxirnately 3.18 crn. The superconducting strands are cornposed of 199NbTi filarnents approxirnately 14.4 J.l.ffi in diarneter, enclosed by a copper sheath approxirnately 3.67 J.l.ffi thick and ernbedded in a Cu0 .9 Ni 0 . 1 rnatrix. The ratio of the constituents in the conductor interior is to be 0.36 parts Cu0 .9 Ni0 . 1 to 1.5 parts Cu to 1.0 parts NbTi. The filarnents are uniforrnly dispersed in this interior region of the conductor. In forrning the conductor, a CuNi sheath is used to hold this rnixed rnatrix together. The resulting overall ratios of Cu0 .9 Ni 0 . 1 : Cu: NbTi therefore becorne 1: 1.5: 1. To reduce the ac loss, the conductor filarnents are twisted approxirnately one turn per 3.18 rnrn. The superconducting strands are insulated frorn each other by an approxirnately 25.4-J.~.rn-thick coating of polyarnide-imide (Westinghouse Omega®) insulation. Table II presents a detailed surnrnary of the conductor characteristics. AC LOSS CALCULATIONS The 400-kJ energy storage coil was designed to have a loss of less than 0.3% of the stored energy for a central coil in a ten-coil stack during a 30-s charge and an 0. 7 -rns discharge. This section presents the loss calculations perforrned to deterrnine the coil performance. The calculation of the ac lasses have been rnade using what are generally regarded as the rnost accurate forrnulas available at present CJ. Table III presents the surnrnary of the loss calculations. 84 S. K. SiJiah, C. J. Heyne, D. T. Hackwortb, M. A. JIUIOCko, P. W. Eckels, and J. H. Mnrphy Table lU. Summary of Loss Calcolations (Losses in Joules) Central coil of 10-coil stack Isolated coil Superconducting Normal Superconducting Normal 0 70.0 462.4 436.7 0 70.0 462.4 0 77.5 381.1 463.7 0 77.5 381.1 Dummy strands Eddy current--charging Eddy current--discharge 0 0.3 0 0.3 0 0.3 0 0.3 Mandrei Eddy current--charging Eddy current--discharge 0 99.8 0 99.8 0 129.5 0 129.5 1069.2 450K 0.238 952.2 450K 0.212 1052.1 400K 0.263 Type of discharge Superconducting strands Eddy current--charging Hysteresis--charging Eddy current--discharge Hysteresis-discharge J2 R --discharge Totallosses Stored energy % Loss of stored energy 319.7 319.7 908.1 400K 0.227 MECHANICAL DESIGN The structure design is based upon the 300-kJ coil [2 ] that has been successfully tested at LASL when subjected to an ohmic-heating cycle eJ. The structural design of the coil was governed by the following design criteria: 1. The entire structure was to be nonconducting. 2. Liquid helium coolant was to be in direct contact with all superconducting strands throughout the coil and Iead area. 3. Combined primary stress intensity should not exceed two-thirds of the material yield stress, or 40% of the ultimate, at operating temperature. 4. The structure must withstand the loading conditions associated with a 10-coil stacked assembly. The main coil structure consists of an inner fiberglass composite tube. This tube also acts as the first-layer coil former and establishes the concentricity of the remaining layers. The superconducting wire is wound into a helical groove in the coil former. Theseturnsare separated by teeth that provide the necessary spacing of the turns and provide structural restraint at axial forces generated in the turns. Cooling ducts are machined in the coil former axially below the turns to provide helium coolant to the first layer. The cooling ducts also provide a channel for helium gas, generated during the pulse cycle, the escape ftom the layer to the gas collection chamber at the top of the assembly. Each wound layer was over-wrapped with resin-coated fiberglass fibers and then cured, in order to contain the turns and to restrain the superconductor against radial forces developed when the coil is energized. To provide adequate layer-to-layer insulation, 11layers of 0.127-mm-thick Kapton polymide film were applied to the outside surface. The remaining two layers of the coil are essentially the same as the first layer, with the exception that insulation was not required on the outside surface of the coil Design of a 400-kJ Pulsed Energy Storage Coil 85 assembly. However, an additionallayer of Mylar was wrapped on the outside of the structure to provide a protective coating for the assembly. The whole coil assembly is then heated for 16 hr at 150°C. This treatmenteures the resin-coated fiberglass wrap on the outside of each former. The material selected for the cylindrical formerswas G-10 Micarta® filament wound material [ 2 ]. The matrix system used for this grade was selected for its crack resistance when subjected to a sudden immersion in liquid nitrogen. The maximum stresses in the former conductor core at operating temperature are -31.3 and 211.4 MPa, respectively. This is weil within the design maximum values for the former and conductor core. The pressure between the conductor and former, based on thesetangential stresses, was calculated tobe 5.71 MPa. Energizing the coil results in radial and axial electromagnetic forces. The resulting hoop stress in the conductor core and the glass tape was calculated as follows. The loading of the system is based on the Ioad sharing between the cured glass tape and the conductor core as a function of the areas and elastic modulus of each. The calculations are based on the assumption that both, glass tape and conductor core, deftect an equal amount due to the radial force. Therefore, 81 = P1L1 A1E1' P2L2 82 = A 2E2 ( 1) (2) 8 = 81 = 82 L = L1 = L2 (3) Therefore, A1E1 (4) A2E2 (5) P = P1 +P2 Using (1) and the fraction of Ioad sharing, the stresses on the conductor core and glass tape can be calculated. Table IV presents the corresponding stresses on the stainless steel conductor core and the cured glass tape. The axial forces are a maximum at the ends of the coil and are directed inward. The 0.508-cm-thick tooth between conductor winding slots must be capable of supporting this Ioad without excessive stresses. Since each former cylinder has 100 0.635-cm-wide axial slots cut along its length, the slot tooth must support an increased Ioad over the magnetic forces. Axialloading will be assumed applied along Table IV. Conductor Core and Cured Glass Tape Stresses Due to Radial Field Forces (Stresses in MPa) Layer No. 1 Layer No. 2 Layer No. 3 Load condition Core Tape Core Tape Core Tape Single coil, 25 kA Single coil, 35 kA Ten-coil stack, 25 kA Ten-coil stack, 35 kA Thermal interference stress 148.6 291.3 173.4 339.8 211.4 28.7 24.5 56.2 48.1 81.4 159.4 112.2 220.0 102.1 18.4 13.4 36.3 26.3 15.7 30.7 37.6 30.7 102.1 6.2 2.6 5.1 5.1 86 S. K. Singh, C. J. Heyne, D. T. Hackwortb, M. A. JIIIIOCko, P. W. Eckels, and J. H. Murphy Table V. Maximum Conductor Loading and Peak Bending and Shear Stress on Slot Teeth Resulting from Axial Loading for Layer No. 1 Loading Stress, MPA Load condition Maximum cond.load, N/cm Total coil Ioad, kN Peak bending stress Maximum shear stress Single coil, 25 kA Single coil, 35 kA Ten-coil stack, 25 kA Ten-coil stack, 35 kA 333 652 350 686 744 1458 851 1669 28.9 56.6 30.4 59.6 10.9 21.5 11.5 22.6 the centerline of the conductor. The large fillet radii of 0.953 cm is used to strengthen the teeth, but if beam theory is applied, the bending stress will be Mc ub = 1 (Px)(h/2) = bh 3 112 6px = (6) h2 Location of the peak bending stress can be determined by plotting the value of x/ h 2 along the length. These values of peak stress are tabulated in Table V at 40% along the tooth length. Designstrengthof the former is 38.6 MPa (0.4uultimate) in the direction transverse to the filaments. Comparison of this value to the values in Table V indicates that at the standard operating condition of 25 kA the tooth is adequately designed with a 0.953-cm fillet. At the extended operating condition of 35 kA the tooth is still within the Iimits if it is assumed that the design strength can be increased to ~ of Uu)timate or 64.7 MPa. The axialload will be less on the teeth toward the central region of the coil, but the Ioad will accumulate in the coil formers to a peak Ioad as shown in Table V. Calculation of the cylinder stress values at 25-kA single-coil Operation gives the stress as 12.4 MPa. For the single-coil extended operation of 35 kA the axial compressive stress on the former is 24.4 MPa. Both of these stress values are weil within the design strength of the former material. THERMOHYDRAULIC ANALYSIS The 400-kJ coil is intended to demoostrate feasibility of a lower-loss, advanceddesign conductor using proven cooling techniques of the 300-kJ coil In appearance, the 400-kJ coillooks very much like its predecessor, the 300-kJ coil, but there is one major difference in the thermal design parameters. The 400-kJ coil thermal stability is reduced from that of the 300-kJ coil. The reasons for the reduced thermal stability are mandated by low-loss requirement and existing conductor manufacturing technology limitations. The cooling channels and interconnecting ducts are designed to be of such dimensions and inclinations that the vapor escapes completely from the coil in the minimum operation cycle time. With escape of vapor assured, cooling during the cycle is accomplished by the evaporation of liquid helium. The operational cycle Iimits of the coil are summarized in Table VI. Note that the coil may become resistive during discharge but must retum to the superconducting state during the rest interval for recharge. Table 111 summarizes the Iosses or e]. 87 Design of a 400-kJ Pulsed Energy Storage CoU Table VI. 400-kJ Coil Cyde Summary Process operation line 0-1 1-2 2-3 3-0 Charge to 25 kA Hold at 25 kA Discharge to 0 A Rest Total period Conductor state SC* SC SC/resistive SC/resistive Minimum interval, s Maximum interval, s 30 3 7 X 10-4 267 300 300 300 1 299 900 * Superconducting. internal heat generated in the conductor during each process. The values given are peak values but are presumed spatially and temporarily constant and enter the thermodynamic states trajectory equations as average values. During coil charging (process 0-1), a hysteresis loss of 10.4 mJI cm3 of conductor is generated in the NbTi filaments, producing a maximum steady-state temperature rise in the filament of approximately 10-8 K, a rise in the conductor strand of 6 x 10-7 K, and a heliumfilm rise of 2 x 10-5 K. In the stainless steel mandrei the Iosses and temperature rise during charging are also negligible. Thus the process 0-1 is essentially an isothermal process. During the hold period (process 1-2), the ripple in the power supply is the only source of loss. The loss due to the ripple in the power supply is 0.2 m W I cm3 averaged over 3-s hold periods. This can also be considered an isothermal process. During discharge, heat is generated within the conductor which is sufficient to induce a transition to the resistive state. Figure 2 shows the computed temperature rise culminating in a final temperature of 16.8 K. In this computation the mode of heat transfer is presumed to be nucleate boiling with a maximum surface fiux at 16.8 K of 0.6 W lcm2 • Some question could exist as to whether nucleate boiling can be sustained within the conductor, but the final average volume fraction of vapor within the conductor not including the helium contained in the risers is 20% and the vapor quality is 3.3%. In the short discharge transient, nucleate boiling is expected to exist. r---------------------------------------, DISCHARGE INTERVAL ~ 18 16 14 "". 12 4 0 L-~~--~~~--~~~--~~~--~~~~ 0 678 TIME, MSEC 101112131415 Fig. 2. Thermal cycle of 400-kJ coil. 88 S. K. smp, C. J. Heyae, D. T.llaebrortll, M. A. JMOCko, P. W. Eckels, ud J. H. Marphy Entering the rest period of 265 s (process 3-0), the conductor contains 20% volume fraction of vapor and is at 16.8 K. No heat generation continues, but the conductor must recover to the superconducting state. During process 3-0, reduced heat exchange is expected to occur within the cable confines. Nucleate boiling is expected to continue in the relatively large open regions of the bubble riser channels. Thus heat will be conducted axially along the conductor to the boiling patch. For a riser channel pitch of 1.6 cm the effective length is 1.0 mm. The axial Biot number of 0.026 indicates that axial thermal impedance can be neglected. Figure 2 shows the recovery transient and predicts recovery of the superconducting state in about 12 ms. Thus the temperature of thermodynamic state point 0 is achieved in 20 ms after discharge, but the coil has not fully returned to its original thermodynamic state because of the 20% vapor quality remaining in the conductor. It must be shown, then, that vapor bubbles can escape from the most distant points in the coil during the rest interval. The bubble velocity is computed for a constant area channel with the expectation that the technique closely approximates the actual rise velocity. In a fluid such as water with relatively high surface tension a variable area riser such as exists in the 400-kJ coilleads to an uneven bubble motion that can significantly slow the vertical rise. In helium, the low-surface tension Ieads to bubble breakup and suppresses the uneven motion regime. The characteristic channel dimension used in the computation is the minimum that exists (6.1 cm) even though bubble velocity correlates best with maximum channel dimension. The time for bubble escape is approximately 14 s, and the initial thermodynamic state is recovered in 34 s after initiation of the discharge. ACKNOWLEDGMENTS The authors gratefully acknowledge the contributions and helpful suggestions of J. D. Rogers and P. Thullen of the Los Alamos Scientific Laboratory. NOTATION A 1 = area of conductor core, 0.204 A 2 =total area of glass tape, 0.88 cm 2 b = tooth width c = extreme fiber distance from neutral axis E 1 = elastic modulus of conductor core, 200 GPa E 2 = elastic modulus of glass tape, 33 GPa h = thickness of tooth (variable) I = cross-section moment of inertia L = length of components M = bending moment P = applied Ioad p = PI b = applied Ioad per unit width n = distance from point of loading to location of stressed action 8 = deßection due to radial force cm 2 REFERENCES 1. C. J. Heyne, D. T. Hackworth, S. K. Singh, M. A. Janocko, P. W. Eckels, and J. H. Murphy, "Design and Fabrication of 400-kJ Superconducting Pulsed Energy Storage Coil," Final Report, Westinghouse Electric Corporation, LASL Contract No. X67-6665 (1979). Design of a 400-kJ Pulsed Energy Storage Coil 89 2. E. Mullan, D. W. Deis, P. W. Eckels, H. E. Haller, 111, M. A. Janocko, S. A. Karpathy, D. C. Litz, C. J. Mole, P. Reichner, Z. N. Sanjana, and M. S. Walker, "Design and Fabrication of 300-kJ Superconducting Energy Storage Coil," Westinghouse Final Report E. M. 5077, LASL Contract No. XN4-32767 -3 (July 1977). 3. F. W. Grover, Inductance Calculations, Dover, New York (1946). 4. T. F. Yang, in Proc. 5th Intern. Conference on Magnet Technology, MT-5 Frascati, Italy ( 197 5), p. 203. 5. P. Thullen, J. D. G. Lindsay, D. M. Weldon, and H. F. Vogel, IEEE Trans. Magn. Mag-15(1):538 (1979). B-4 OPERATING CHARACTERISTICS OF A 1.5-MJ PULSED SUPERCONDUCTING COIL* S. H. Kim, S.-T. Wang, and M. Lieberg Argonne National Laboratory Argonne, Illinois INTRODUCfiON The ohrnie heating coils of tokamak fusion reaetors require a stored energy on the order of 1 GJ, a ramping rate of -9 T/s, and a peak operation eurrent of 50 to approximately 100 kA. Owing to power balanee requirements of the reaetors at these sizes, only supercondueting eoils are expeeted to be economieal. For the development of the ohrnie heating eoils, a eryostable pulsed supereondueting coil has been eonstrueted and tested at Argonne National Laboratory. The eoil has a stored energy of 1.5 MJ and a peak field of 4.5 T with a peak operation eurrent of 11.2 kA. The eoil was tested with both de and pulsed eurrents. FABRICATION OF THE CABLE The design of the cable for the 1.5-MJ coil is based on detailed cryostability The cable was fabricated by studies of basie cables and 5-kJ model coils Supercon, lne., to Argonne National Laboratory specifieations by twisting 24 basie eables around an insulated stainless steel strip with a twist piteh of 22.5 cm. A close-up of the eable eross-seetion is shown in Fig. 1. The basie cable is made by twisting three seven-strand conductors (triplex cable) with a twist piteh of 2.2 cm. The seven-strand conduetors are made of six OFHC copper wires twisted around a superconducting center conductor and soldered with Staybrite. Since the requirements of low ae Iosses and cryostability confiict with eaeh other, the basie principle chosen is to achieve cryostability within the basic cable. To restriet ac coupling among the 24 triplex cables in the final cable, only limited current sharing among the triplex is allowed by coating a thin insulating film around the seven-strand conductors. The critical current of a short sample of the basic cable is 405 A at 5 T. Each superconducting strand has a diameter of 0.051 cm and contains 2041 6-JLm filaments with a twist pitch of 1.27 cm. The copper-to-superconductor ratio for each strand is 1.8. The final cable is compressed during the cabling by heavy rolls from four sides. This is required to minimize mechanical perturbations of the basic conductors during pulsing of the 1.5-MJ coil. The compression did not darnage the insulation between eJ * Work supported by the U. S. Department of Energy. 90 Operating Characteristics of a 1.5-MJ Pulsed Superconducting Coil 91 Fig. 1. Cross section of ac cable. the stainless steel strip and the 24 triplex cables. However, owing to the deformation of the soft soldering in the seven-strand conductor, about 5% degradation of the recovery current in the triplex has been observed. The finished cable has a width of 3.78 cm and a thickness of 0.74 cm. The first 25-m-long cable was produced as a test of the cabling technique. The total cable length fabricated for the coil was about 590 m long. 1.5-MJ COIL FABRICATION Coil winding with a spongy cable is a rather interesting experience. The coil is composed of 18 helicallayers with an average of 14.3 turns in each layer. Turn-toturn insulations are provided by two layers of 0.02-cm-thick glass-cioth tape and two layers of 0.01-cm-thick Mylar tape. The winding layers for the first to tenth layers are separated by 0.48-cm-thick and 0.64-cm-wide G-10 strips. In the low-field region of the eleventh to eighteenth layer, the thickness of the strips is reduced to 0.32 cm. Characteristics of the coil are listed in Table I. Figure 2 provides a close inspection of the coil winding in the thirteenth layer. Spaces between the strips provide 0.64-cm-wide cooling channels in the vertical direction. The G-10 bobbin has cooling channels in the radial direction. During the winding, the tension in the cable was increased gradually from 225 to 450 kg to provide a constant radial pressure in the coil. Totallength of the cable used for the coil is about 510 m. Nine potential taps have been installed in the coil to study possible conductor motion and six thermocouples to monitor temperature variation during the test of the coil. When the coil is charged, the average hoop stress in the cable is about 310 MN/m 2 • Part of the stress can be sustained by the stainless steel strip in the cable itself. To support the excess stress, 16 bands are placed outside the coil. Each band is fabricated of 30 layers of 0.025-cm-thick by 3.2-cm-wide fiberglass cloth. EXPERIMENTAL ARRANGEMENT FOR TUE COIL TESTS The 1.5-MJ coil is assembled for tests as shown in Fig. 3. Details of the plastic Most of the cryostat developed for the coil have been reported elsewhere materials used for the arrangement are nonmetallic to avoid eddy current Iosses during the pulsing of the coil. The topflange is constructed of 5.7-cm-thick Micarta plate. The coil is suspended to the top flange using eight 0.64-cm-diameter stainless e]. 92 S. H. Kim, S.-T. Wang, and M. Lieberg Table I. Characteristics of Pulsed Superconducting Coil Central field, T Peak field, T Operation current, kA Inductance, mH CoiiiD, cm Coil OD, cm Axiallength, cm Number of layers Total number of turns Cryostable recovery heat flux, WI cm 2 Layer-to-layer spacing, cm Average current density, Alcm 2 Cable cross section, cm Cable length, m Total ampere-meters, A-m Maximum radial magnetic pressure, MPa Maximum axial magnetic pressure, MPa Maximum dBidT, Tls Maximum dll dT, kAis Charging voltage, V Hysteresis loss in the filaments, kJ I cycle Eddy current loss in the matrix at 9 Tls, kJicycle AC losseslstored energy at 9 T ls, % Eddy current loss in the stainless steel at 9 Tls, Jlcycle Heat flux due to ac Iosses at 9 Tls, mW lcm 2 4.2 4.5 11 24 41.6 81.0 58.1 18 258 0.35 0.48 (1st-10th layer) 0.32 (11th-18th layer) 2290 (1st-10th layer) 2685 (11th-18th layer) 3.78 X 0.74 510 5.8 X 106 83 28 11 27 650 -0.1 2.65 -0.1 60 -10 steel rods. The bottom of the coil is further supported by a 2.54-cm-thick Micarta plate. Thermalradiation shields consisting of a 10-cm-thick piece of Styrofoam and eight layers of aluminum foil are attached to the bottom of the top ftange. The two coil terminals are brought to the top of the coil by gradually changing the winding angles. After removing the thin insulation in the basic cable, each terminal is soft soldered to eight copper-stabilized monolithic superconductors, 1 cm wide and 0.2 cm thick. This design allows a change in the directions of the terminals and makes the terminals mechanically solid. The terminals are connected to the bottom tips of the vapor-cooled current Ieads. The Ieads purchased from American Magnetics Inc., have a current capacity of up to 15 kA dc. Heat leaks of the Ieads are about 22 W without current and 30 W with 12 kA dc. During the test of the coil the liquid helium Ievel should be maintained between the bottom tips of the Ieads and the top of the coil. The distance between them is about 23 cm; this corresponds to a helium boil-off energy of about 380 kJ. With this energy the coil can be pulsed about 140 times with a pulsing rate of 9 T /s. DC SHARING The 1.5-MJ coil was fi.rst charged to the critical current of the short sample cable (11.2 kA) by a 5-V, 12-kA dc power supply. During the fi.rst charge of the coil no major conductor motion or mechanical perturbation was observed. Figure 4 shows the critical current and the Ioad line of the coil. The critical current of the cable was determined from measurements of the critical current of short sample triplex cable. Operating Characteristics of a 1.5-MJ Pulsed Superconducting Coil Fig. 2. Thirteenth layer of 1.5-MJ coil winding. Fig. 3. Experimentalarrangement for the 1.5-MJ coil. 93 94 S. H. Kim, S.-T. Wang, and M. Lieberg >- "' I, kA Fig. 4. Critical current of short sample cable and Ioad line of the 1.5-MJ coil. To demoostrate the cryostability of the coil, it was charged beyond the critical current up to 11.75 kA (from point A to pointBin Fig. 4). A bridge circuit was used to detect when parts of the coil in the high-field region became normal. Beyond the critical current, unbalanced valtage of the bridge increased gradually; this indicated a stable current sharing between the superconducting strands and the copper stabilizer. The current-sharing section, with a resistive valtage of 2 mV, was estimated to be about 1.5 m lang. Charging the coil to point A without developing a resistive valtage is a significant result. This means that the cable is fully transposed and the current-carrying capacity of each of the 24 basic cables is equal without any degradation. PULSED-CURRENT TESTS Single Pulsing. After the current-sharing test, the coil was pulsed with a 7 -MW (650 V at 10.9 kA) power supply. A summary of the pulsed-current characteristics is shown in Table I. The coil was charged to a 4.4-T peak field in 0.4 sec and discharged to zero in 0.6 s with a maximum ramping rate of 11 T /s. The off-time between pulses was 10 s. The terminal valtage of the coil, V ooih can be written as (1) where L is the inductance of the coil and V1oss is the valtage associated with the energy Iosses in the coil during pulsing. Figure 5 is a set of recordings for a typical pulsing test. In this figure the peak current is 10 kA with a pulsing time of 1 s. The left side of the figure is an expansion of the recording, and the right side shows a continuous pulsing between off-time of 10 s. The current variation dl/ dt is close to a triangular waveform (Fig. Sa) with a maximum rate of 27 kA/s. This rate is not limited by the coil performance but by the power supply used. The terminal valtage of the coil, V ooii (Fig. Sb), is balanced with an inductive valtage taken from a mutual Operating Characteristics of a 1.5-MJ Pulsed Superconducting Coil 95 Fig. 5. Recording of a single-pulsing test. Legend: (a) current waveform; (b) coil terminal voltage; (c) loss voltage; and (d) integrated loss voltage. inductor placed outside of the coil. The resulting loss voltage, V coih is shown in Fig. Sc. Double Pulsing. After more than 3000 single pulses the coil was tested with double triangular waveform pulses as shown in Fig. 6. The double-pulsing modewas used to simulate the full ftux swing of ohmic-heating coils. The peak current during the pulsing was 10.6 kA with a central field of 4.0 T. The current waveform is shown in Fig. 6a. The full period of the double pulsing was 9.5 s with an off-time of 6.9 s. The charging and discharging times were about 0.64 s, respectively, in each pulse. The double pulse at the left side of the figure is an expansion of the pulses at the right side of the figure. Potentialleads are taped to several layers of the coil. The potentials between layers 2 and 4 and between layers 12 and 14 are shown in Figs. 6b and 6d, respectively. After 570 double-pulsing cycles no visible differences between the two voltage waveforms have been observed. AC LOSSES The ac Iosses of the coil were determined from helium boil-off during the pulsing and by the electronic integrator method 3 ]. Alternating current Iosses as a function e· 96 S. H. Kim, S.-T. Wang, and M. Lieberg ~ ~ !II~ l l l lllilllllll!i I : ;;i !tl:; I:il 1:;t11i:: llil!l~~~~~:! :;, I: :II:;.IIIJIJJIIIII!IIIItlllllll!fl!!lll!!!l!! ll!!l\1!\ll!llll!lli,lll!;l:I: :!::11;1~! II!/: ;~~~;;;1 ilh 1!11" um 'i : ,, I' 11111 tll[l II II !~!I ~ll :: :: Ji:;: 1:: :lli~ 1::: : !:~! ~il:: 1:1 • :: : : :· :': ,,, '~'" 11111 '';''111H'Hl!l1JIIi1J dtlhHJ!lU lllHl!illlUillifl::umHmi!!IH'I !'!ll!!U:I:H!HI!'!Il!UJ!'I111 !llllll lll!llliüJ!rll'lllll' U!IJll!n!IIJiill!rllll·:l'lll!%"h'·, 111111111 1111 H:l!ll::u::wm~ l:.l:l!IJJ:! l,l' p, lt ,, ,' ,, 111. IIIIIJ!I'' , '!· l; l !11!1Nil!' ' W!!H !J!J1 lllitllillllllllih,i!llltf,litt,iiiii,J:iHm:,idi::·l;lii:HI'i:H'!!!ll:jl.l'li:J:::t:ntlm]!:! 1 Iu "11: :• nl •; !I!Wll!l! lliiLIIfJillllllllln!Jiillllllllltllii lnlliUI!m•miHhiE!Il,l.i, 1tmh:dl iilWH!!HI!! 1!1! I!U ll~llUW! 11111 'l!'llllll! 'J 11 •11ll Jhilullll' 11.111 IJTII , :J'I!:'nlliiUIIIUIUillumllmll!~!liiiiii!JJml!'llliUUJillmiltlllllliiiiiillllmnnUIIlllllliiii!IIHIIIIIIIHU,Jlilniiillilllli:ll!IH,.!i !l:l jl::·:~·· J::l''!'Jl'l "U ' l!II!JI'Jll• J" JrnriJIIIIIIJ mln,u,ll.llllilllhllh ;::ll!:lilllil'llll 111 lllll H!!H:!JIP!1l1!11!!!!1!11 1PII11!1 1111U!!!!ü!l!l1!n llllllfllflllllllllllllllilllliilllllli!lll•llillllilil iillnlli!lu,lml!l;li•ln•l•illii:l:ll :a:: !!:!l'lll' -111'11111'111' l!!l'li~llllil lllllll!lllllllillllilllllll• ,,,.;:,lillllllhmllr~MIIilliiilnll:mhilil!llllllllllllllllll~l!ill:l'il'l!"f!U,!IIl!IH!l!l!llll!ll'llll!llll!lllll!,!!lll!lll!ll!l!llllllll!flllll'll'limllullllllill.iihiniiii:.rr:ll,l::r::ll:i:llll:a. r:r:'!lildllhllll!llliH!!!I'I!IP.I!''!l!' ''l"''llltillllllllllilllillili•lliuliilliilllilillllllllllllllllill;h!l,,r;; .. i!,;u;m::lll!l.tmlll::lljll'i'llllllilll'i'lllllllllllll!ll!lllllljlllll!ll!l!lll'l'llllll'l'!l!"!!'l'lilluilllllllhill•llliilli:lin:r,r,·riillllilll:lllliliHlilrtl'i'lm::r:r:· ''""IIIIUIIIII!!II!'III!!IIIIIIIIIililllli!llltilnllillllllillllllililll!lnllll!!lilll•l,lllil'lililllllllll II , I II lllililllllllll!lllllllll!l'l:!l!i:I!!Jil!l'l 1 ! 1 1! 1 1 1 llll!l!llllilil!illiilliinliilliililii,lliillliiiliillliil~li11 il!i,inll' Fig. 6. Recording of a double-pulsing test. Legend: (a) current waveform; (b) valtage of 2nd-4th layer of coil (2 V /div.); (c) loss voltage of coil (0.2 V /div.); and (d) voltage of 12th-14th layer (5 V /div.). of (dB/ dt) 2 are shown in Fig. 7. Data points shown as circles and triangles are obtained from the tests with single pulses, and data points shown as solid rectangles are obtained from the double-pulsing test. Eddy current Iosses in the copper stabilizer can be expressed as (4 ] (2) where Veu is the volume of copper, lp is the twist pitch length, R is the radius of superconducting wire, p 1 is the effective transverse resistivity, and R, is the conductor radius including copper stabilizer. The linear variation of the ac Iosses as a function of .8 2 in Fig. 7 indicates that most of the Iosses are a result of the eddy current in the copper. The insulation on the surface of the six-strand wireisthin enough to have limited This allows a certain degree of ac current sharing among the triplex cable coupling among the 24 triplex cable of the 15-MJ coil. If the coupling is assumed to be limited within the six-strand wire, from the experimental data of Fig. 7 and (2) the effective resistivity is found to be about 2 x 10- 10 0-m, which is somewhat lower than expected. This is an indication that some portion of the ac Iosses come from the ac coupling among the triplex cables and the six-strand wires. From the above results one can conclude that the thin insulation of the 1.5-MJ coil cable has an Optimum thickness to compromise the current sharing and low ac lasses. eJ. Operating Charac:teristics of a 1.5-MJ Pulsed Supercouducting Coil 2.5 I lo/ / 2.01- 31: oo / VI !S --' :;: 1.0 1- / y 0 / yo'Y - 6 - /0 ~% Q:)Cö 0 0 6 ~o/ 6 1.51- 0.5 1- - oo/ o':/ r 10 97 0 I 20 30 40 I I I 50 60 70 80 (d8tdt) 2 , ( Ttd Fig. 7. Ac Iosses vs. (dB/dt) 2 . Data points of circles and triangles represent single-pulsing tests while solid rectangles .represent double-pulsing tests. PULSING EFFECf ON THE CRYOSTABILITY After the pulsing tests of the coil, another dc current test has been conducted using a 50-kA, 5-V dc power supply. This testwas designed to investigate the pulsing effect on the cryostability of the coil. The coil was charged with a charging valtage of 0. 7 V. No significant change in the critical current was observed. The coil remained in a current-sharing state up to 1 kA above the critical current, and recovered to the superconducting state by reducing the current. The coil was quenched when the current was further increased. The amount of the energy released in the quench estimated from the blow out of the helium was approximately 0.5 MJ. In the subsequent charging, discharging, and quench, no significant changes in the characteristics of the coil have been observed. SUMMARY 1t has been demonstrated for the first time that a relatively large cryostable superconducting coil can be pulsed with relatively low ac Iosses at high ramping rates. The low ac Iosses with limited current sharing have been achieved by insulating the surface of the basic cable with a thin organic film. The ac Iosses at 9 T/s are about 0.1% of the stored energy in the coil. After more than 4000 pulsing cycles, no changes in the pulsing characteristics and cryostability of the coil have been observed. The quench current of the coil was approximately 1 kA higher than the critical current of the coil. REFERENCES 1. S.-T. Wang,S. H. Kim, L. R. Turner,K. M. Thompson, W.F. Praeg,C.I. Krieger,and R.L. Kustom,in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 255. 2. S. H. Kim, S.-T. Wang, W. F. Praeg, C.l. Krieger, and M. Lieberg, IEEE Trans. Magn. Mag-15:840 (1979). 3. S. H. Kim and S.-T. Wang, "Fusion Power Program Quarterly Progress Report," Argonne National Laboratory, ANL/FPP-78-4 (1978). 4. W. J. Carr, Jr., J. Appl. Phys. 45:929 (1974). B-5 3-MJ MAGNET FOR SUPERCONDUCI1VE ENERGY STORAGE T. Shintomi, M. Masuda, H. Sato, and K. Asaji National Labaratory for High Energy Physics Ibaraki, Japan INTRODUCfiON Superconductive energy storage requires different kinds of magnets depending on its applications. Load leveling of electric power requires a cryogenically stabilized magnet, similar to the dc magnets used for bubble chambers. The stored energy for such magnets is 10 12 to approximately 10 13 J which is larger than the stored energy of magnets for other uses. One of the most important problems is how to reduce Iosses at the interface between the superconductive coil and the power grid. Another type of magnet utilizing pulsed superconductive energy storage is that used for accelerators and fusion reactors. This magnet should be also cryostable since it is required to transfer a huge amount of energy in a short period of time, frequently on the order of 1 sec. lt is necessary, however, for this application to use a nonstabilized superconductor in order to suppress ac Iosses which result from the pulsed operation. The dewar vessel for a pulsed magnet requires more stringent specifications because the eddy current which results from a fast change in the magnetic field significantly increases the heat Ioad at liquid helium temperature. A plastic dewar must be used to prevent eddy currents even if it is prone to helium leaks through the plastic wall. Aseries of investigations C-3 ] by the present authors that uses 100-kJ coils has demonstrated that energy storage for an extended period of time does not require a superconductive switch to maintain the persistent current and energy transfer using a thyristor inverter with capacitors is one of the best methods for pulsed power Ioads, especially for large amounts of pulsed energy storage. A practical evaluation of such energy storage, however, requires data using coils much larger than 100 kJ. A 3-MJ energy storage magnet is thus the next step toward a full scale magnet. The 3-MJ coil was constructed to ascertain the following information: 1. 2. The efficiency of the coil when used as an energy storage device for Ioad leveling of an electric power system measured by energy transfer to apower line utilizing thyristor bridges. The energy transfer between the 3-MJ coil and two 100-kJ coils using the thyristor bridges proposed by the group at the University of Wisconsin [4 ]. (This will provide the design concepts for the pulsed energy storage associated with fusion reactors, accelerators and other pulsed power Ioads.) 98 99 3-MJ Magnet for Superconductive Energy Storage Table I. Design Parameters Coil specification Stored energy, MJ lnductance, H Mean radius, cm Length, cm Winding thickness, cm Number of turns Nominal current, A Central field, T Maximum field at wire, T Weight of wire, kg Weight of bobbin, kg 3 5.4 60 29.2 8.0 1773 1050 1.9 5.0 280 310 Wire specification Size, mm Matrix ratio (Cu: Nb Ti) Filament diameter, J.Lmrp Number of filaments per wire Wire twist pitch, mm Critical current at 5 T, A B A 1.5 X 3.0 3.0 24.5 2257 30 1500 1.5 X 3.0 3.0 61.5 361 40 1600 3. The self-protection of the coil upon intentionally quenching one section of the divided coil. The design and construction of the 3-MJ superconductive energy storage system and also the preliminary experimental operation are reported in this paper. COIL Specifications To obtain high efficiency as an energy storage magnet, the coil must have a large inductance, particularly for load-leveling applications. Judging from previous studies with the 100-kJ unit, the efficiency of the 3-MJ unit was estimated to be approximately 70% for repetition times on the order of 10 min exclusive of cryogenic and mechanicallosses. The efficiency is given by [5 ] 71 = 1 - ~ aRPT- 2n Ll v(;J 1/2 T (1) where the factor, a, is inversely proportional to the inductance of the coil. This implies that with a larger inductance, a larger system efficiency should be possible. The rated current in this study was fixed at 1050 A to obtain the above target efficiency. The coil thickness was designed so that the field at the coil end did not exceed that at the inner surface of the coil in the medium plane. Designparameters of the coil are listed in Table I. Superconductors A and B were fabricated by Kobe Steel and Hitachi, respectively. Structure The salient points of the design are that the coil is divided axially into three sections and that the superconductors are reinforced by stainless steel tapes as shown in Figs. 1 and 2, respectively. The segmentation functions not only to Iimit the pressure exerted on the windings caused by axially compressive forces but also to 100 T. Sbintomi, M. Masuda, H. Sato, and K. Asaji IO.Scm 120cm'l> Fig. 1. Cross section of the 3-MJ coil. Fig. 2. Schematic view of the winding. protect the coil. When one of the coil sections quenches, its energy is partly transferred to adjacent sections in a superconducting state through mutual coupling and partly to a power line by means of inverter operation of thyristor bridges as shown in Fig. 3. This mode enables the quenched section to discharge its energy faster than by the usual protection modes. To achieve such a large inductance the coil was fabricated with a large number of turns of a relatively fine wire, in contrast with other fast-pulsed magnets [6 ' 7 ]. At the same time, the coil bad to have such characteristics as high strength against magnetic forces and excellent cooling for pulsed operation. The superconductors were supported against the radial expansion force by stainless steel tapes which were wound together with the superconductors in a bifilar winding. Good cooling was obtained by winding B-stage epoxy impregnated tapes spirally around the superconductors. The wire was covered about 40% with this tape with a thickness of 0.53 mm. Construction The superconductor was coated with a 40-50-~tm thickness of polyimid to provide adequate insulation protection against the test voltage of 4 kV before applying the B-stage epoxy tapes. This extra insulation protects the coil against the high voltages encountered in pulsed operation. Breakdown voltage of the coil was determined tobe over 2 kV. The coil was wound under a tension of 100 N and cured at a temperature of l20°C for 7 hr. The superconductors were reinforced by stainless steel tapes and secured by the epoxy-impregnated tapes. The stainless steel tapes acted not only as reinforcing members but also as a surface for reliable bonding with the epoxy impregnated tapes. The bobbin was constructed of stainless steel and segmented to reduce eddy currents induced by fiux change encountered in pulsed operation. FundaiTIIItltal cct Superconducting state Normal state of one segiTIIItlt Fig. 3. Schematic circuits developed for coil protection. 101 3-MJ Magnet for Superconductive Energy Storage 120cm95 q 1-' I o - lloHI-+IJ""'"""' I~ I Fig. 4. Schematic of the cryostat. CRYOSTAT The cryostat for the 3-MJ systemwas constructed of stainless steel although a plastic Dewar is more suitable for pulsed operation. The cryostat was designed as a scaled-down model of what might be used in a practical superconductive energy storage system, i.e., the coil has a small aspect ratio and a small heat leak. Even so, it is short in height as shown in Fig. 4. In order to reduce the heat leak, the walls of the helium vessel were fabricated of thin stainless steel and reinforced by thick stainless steel members. Cooldown tests were performed on the assembled cryostat. The estimated volume of helium required for cooldown is shown in Table II. During actual cooldown a total of 450 Iiters of liquid helium was transferred into the vessel with a resultant boil-off of 170 Iiters. Heat leak was measured by monitaring the liquid helium Ievel with a superconductive Ievel sensor. The measured heat leak was 3. 7 W. PRELIMINARY TEST OF COIL Excitation of Coil The assembled coil before placement in the cryostat is shown in Fig. 5. The coil was excited by using two thyristor converters in series. Each section of the coil was tested before the entire coil was excited. While one section was under excitation, the others were short circuited with 8-0 resistors. The middle section was excited first and attained a current of 480 A after six quenches. The lower section was then tested and attained a current of 500 A after five quenches. The upper section did not show any quenching characteristics up to a current of 500 A. Table II. Estimate of Liquid Helium Consumption during Initial Cooldown Liquid helium, Iiters Components Bobbin and structure Cooling channel Cooling channel Wire He vessel Materials ss 304 Phenolic resin B-stage epoxy tape Cu+NbTi SS 304L Total Weight, kg max. min. 294 16 7 273 130 720 429 48 21 549 190 1237 32 5 2 41 14 94 101 T. Shintomi, M. Masuda, H. Sato, and K. Asaji Fig. 5. Assembled coil before placement in the cryostat. The entire coil was finally excited, and after six quenches attained a current of 685 A . The coil returned to its superconductive state after some of the quenches when the current was decreased below the quenching current (see Table 111). In some quenches the coil did not recover and about 100 liters of helium were evaporated. Queuehing The quenching which was observed during training may be classified into two modes. The first mode is associated with the generation of a localized normal state in the superconductors characterized by a low propagation speed. As shown in entry # 2 of Table 111, the propagation velocity of the normalzonewas estimated as 0.63 m/s using the derivative of the unbalanced coil voltage of 15mV/sand the substrate resistance per unit length (38 #LÜ/m at 4.2 K). When the normal region appeared, the current decreased because the control function of the thyristor was changed to an inverter operation. The superconductive state recovered at a current of 528 A within about 10 s after quench without excessive evaporation of helium. This suggests that a few turns of the coil went normal but recovered because of the low normal state Table ßl. Observed Propagation Veloeides and Recovery Currents at the Quenches Quench No. Quench current, A #1 #2 #3 #4 625 635 635 655 Propagation velocity, m/ s Recovery current, A 1.1 517 528 460 540 0.63 1.2 1.1 103 3-MJ Magnet for Snperconductive Energy Storage propagation velocity and the rapid initiation of the inversion operation of the thyristor bridges. The second mode of quench was followed by considerable evaporation of helium at a current of 678 A. Later, the coil deformationwas carefully checked and it was found that the upper section bad moved toward the middle section incurring a maximum deformation of 3.3 mm. Judging from the deformation of the flange and the calculated electromagnetic force equation, the axial compressive force was estimated at 170 tons. lt can be shown that when a coil is moved adiabatically over a distance of 1 mm on the average, the entire superconductor could easily increase its temperature sufficiently to exceed the critical temperature in a short time. The time, Tv, needed for the current of the quenched coil to decrease to zero is given by the following equation, assuming the entire coil is in the normal state: L ( 1 +-RJo) Tv =-In Re VI (2) This time for the test coil is calculated to be 22 s assuming the temperature of the entire coil is below 20 K. When the temperature of part of the coil is above 20 K, the calculated value of time is in good agreement with the measured one of 17.0 s. Propagation Velocity of aNormal Zone The propagation velocities of anormal zone were measured for several currents and the results are listed in Table III. They range from 0.63 to 1.2 m/s. Dresner [8 ] has calculated the velocity for the case of constant thermophysical properties, including the effect of current sharing. Stekly's parameter, a, is estimated to lie between 10 and 40; however, there are uncertainties owing to incomplete knowledge of the heat transfer coefficient and the effect of the spirally wound tapes on the superconductor. The data for low propagation velocity of about 1 m/s and the measured recovery currents are consistent with the value of a calculated by Dresner. The currents reported for quenches in the first mode are the critical values at which the normal zones propagate or recover. Operating Efficiency Measurements were made to estimate the operating efficiency of the system. The efficiencies obtained are shown in Fig. 6 for a stored energy Ievel of 0.85 MJ. 1.0 STORED ENERGY O.B5 MJ > u 0 ~ .5 ü Li: "w Fig. 6. Operating efficiency vs. discharging time for the coil. %~--------~5~0----------7,100~-­ DISCHARGING TIME , sec 104 The data can be represented by the following relation: 71 = 1 - 0.0050T (3) The slope of this line is determined by the Iosses in the power supply and Iead wires. The circuit resistance is calculated to be 18 mß if this relation is utilized. This value agreed with the eflective resistances of the power supply and the Iead wires. SUMMARY The 3-MJ superconducting solenoid, which is intended not only for the power leveling but also for the pulsed storage, was constructed by using monolithic multifilament superconductors. The coil was reinforced to withstand the radial expansion force by the stainless steel tapes wound in the bifilar winding. On the other band, the axial segmentation of the coillimited the pressure exerted on the windings caused by the axial compressive force. Measurements of the propagation velocity and the recovery current were performed in a preliminary test. The values obtained are consistent with previous measurements made by Dresner. The coil experienced quenches because of wire movement which immediately followed deformation of the structure. To minimize this movement, the coil structure will be reinforced in the axial direction. ACKNOWLEDGMENTS The authors gratefully acknowledge the contributions of Y. Oda, N. Sato, M. Nakano, Y. Hayashi, and their collaborators of the Tokyo Institute of Technology. NOTATION E, = maximum stored energy of the coil 10 = coil current at quench L = inductance of the coil n = number of thyristors of the power supply in series Re = resistance of the coil at quench RP = effective resistance of the power supply and Iead wires T = time for charging or discharging coil T0 = discharging time at quench V1 = inverter input voltage of the power supply at quench ~V = forward voltage drop of the thyristor a = 2/L 71 = efficiency of energy storage without cryogenic and mechanicallosses REFERENCES 1. M. Masuda, T. Shintomi, S. Matsumoto, H. Sato, and A. Kabe, in Proc. 6th Intern. Conference on Magnet Technology, ALFA, Bratislava, Czechoslovakia (1977), p. 254. 2. M. Masuda, T. Shintomi, H. Sato, and A. Kabe, IEEE Trans. Magn. Mq-15:318 (1979). 3. M. Masuda, T. Shintomi and K. Asaji, in Advances in Ctyogenic Engineering, Vol 25, Plenum Press, New York (1980), p. 61. 4. H. A. Peterson, N. Mohan, W. C. Young, and R. W. Boom, in Energy Storage, Compression, and Switl:hing, Plenum Press, New York (1976), p. 309. 5. M. Masuda and T. Shintomi, Ctyogenics 17:607 (1977). 6. C. J. Mole, D. W. Deis, P. W. Eckels, H. E. Haller, M. A. Janocko, S. A. Karpathy, D. C. Litz, E. Mullan, P. Reichner, Z. N. Sanjana, and M. S. Walker, in Advances in Ctyogenic Engineering, Vol 23, Plenum Press, New York (1978), p. 57. 7. S.-H. Kim, S.-T. Wang, W. F. Praeg, C. I. Krieger, and M. Lieberg, IEEE lrans. Magn. Mac-15:840 (1979). 8. L. Dresner, IEEE lrans. Magn. Mq-15:328 (1979). B-6 CONCEPTUAL DESIGN OF A 20-MJ SUPERCONDUCTING FORCED-COOLED OHMIC-HEATING COIL* S. K. Singh, J. H. Murphy, M. A. Janocko, H. E. Haller, D. C. Litz, and P. W. Eckeis Westinghouse Electric Corporation Pittsburgh, Pennsylvania and J. D. Rogers and P. Thullen Los Alamos Scientific Laboratory Los Alamos, New Mexico INTRODUCfiON Conceptual design of a 20-MJ superconducting coil is described which was developed to demoostrate the feasibility of an ohmic-heating system. The superconductor material was Nb 3 Sn for a 9-T maximum field. Cabled and braided conductors were investigated and the braided conductors were identified as the best alternates because of their high operating current densities and their high porosity C]. The coil was designed to be cryostable for bipolar operation from a +9- to -9-T maximum field within 1 s. The forced-cooled design described in this paper utilizes crossftow cooling. The coil was designed to generate the ftux swing while simultaneously meeting the limitations imposed by cooling, insulation, current density, and stresses in the materials. DESIGN CONSIDERATIONS The superconducting ohmic-heating coil described in this study required an energy store of 20 MJ at a current of 50 kA. The coil structure was to withstand the forces resulting from 60 kA carried by the coil. The required coil inductance was 16 mH, resulting in a terminal voltage of 1600 V for a resistive discharge and 2500 V for a capacitive discharge. The coil was designed for bipolar operation from +9to -9-T maximum field on the conductor, and bipolar half-cycle sinusoidal * Work supported by University of California, Los Alamos Scientific Laboratory, Contract No. L488407C-1. lOS 106 S. K. Siqla et tll. Table I. Coil Design Parameters Energy storage rating at 50 kA, MJ Peak field, T lnductance, mH Insulating rating, k V Type of cooling Type of conductor Superconductor material Coillength, cm Coil diameter, cm Coil bore, cm Number of turns Number of layers 20 9 16 10 Forced Braid Nb3 Sn 139.7 149.4 33.0 240 10 operation from full positive field to full negative field within 1 s. The hold time between field reversal was to range from 10 to 100 sec at full field. The coil wastobe cryostable and not go normal durlog bipolar operation. The superconductor material was Nb3Sn for a peak field of 9 T. A multilayer helically wound coil design was selected because of the stored energy requirement, high operating current density, and the desire for a fully ventilated coil. The resulting design utilized a lattice braided conductor in a ten-layer configuration. The major design parameters are shown in Table I. eJ CONDUCI'OR DESIGN Several conductors were designed for potential application in a forcedcooled ohmic-heating coil. The superconducting material considered was a Nb 3Sn braid. All conductors were designed to be cryostable, i.e., the heat generation rate when all of the current is in the stabilization material is less than the heat removal rate. The copper-to-noncopper ratio and the strand diameter were selected to give an operating current density to critical current density ratio approximately equal to 50%. AC LOSSES The ac Iosses in the ohmic-heating coil conductors have been calculated assuming a nonoptimized superconductor configuration. Table II summarizes these calculations and the reference conductor characteristics. The predominant ac loss, in all cases, is eddy current loss in the copper stabilizer. This loss can be lowered by incorporating resistive webbing. However, because the surface heat ßux during the field pulses is between 2 and 14 times smaller than the surface heat ßux required for cryostability, this loss reduction has been ignored at this time. MECHANICAL DESIGN The structural design and analysis were based on the 300-kJ coil [3] that has been successfully tested at LASL when subjected to an ohmic-heating cycle [4 ]. The pulsed operaring cycle requires minimization of electrically conducting structures to minimize eddy current Iosses during the operating cycle. For this reason fiberreinforced composites were used wherever possible in the coil structure. The structural designwas developed using the following design codes: 1. The maximum strain in the Nb3Sn superconductor should be limited to 0.1%. Conceptual Design of a Superconducting Ohmic-Heating Coil 107 Table II. Reference 50-kA Conductor Characteristics Peak field, T Type of cooling Type of conductor Material Number of strands Bare strand diameter, cm Insulated strand diameter, cm Number of filaments Filament outside diameter, J.Lm Filamentinside diameter, J.Lm Filament twist length, cm 1 to Je ratio Cu-to-non-Cu ratio Packing factor Conductor width, cm Conductor thickness, cm Conductor current density, kA/cm 2 Matrix conductivity, um, mho/m Transverse conductivity, u .L • mho/m Sheath conductivity, u, mho/m Peak ac loss per pulse, J/cm 3 Peak field pulse heat ftux, mW /cm 2 Cryostability recovery heat ftux, mW/cm 2 9 Forced Braid Nb3 Sn 503 0.0965 0.102 6,328 3.5 1.5 1.0 0.500 3.069 0.41 3.50 2.86 5.00 5.9 X 108 1.2 X 109 1.7 X 109 1.66 58.0 448 2. The primary stress intensity in the structure should be less than two-thirds material yield stress or 40% of ultimate, whichever is less. 3. The maximum stress theory will be used for composite structures. The design concept for the coil is shown in Fig. 1. The concept uses teeth to support axial forces and stainless steel bands for radial support. The axial and radial forces vary within the coil cross section. FilAMENT WOJNO KEYS C()IJOUCTOR SUPPORT GROOVE AXIAL STRUCTURE HELIUM COOlANT CHANNEL Fig. 1. Cross section of a layered wound coil. 108 S. K. Slngb et al. The structural design considers three areas, the tooth thickness, the banding thickness, and the former thickness. The tooth thickness calculations are based on a cantilever beam with uniform loading in the axial direction. The tooth width can be decreased from its maximum value at the coil ends toward the center as the axial force decreases. The banding concept uses an overwind of metal band on the conductor outside surface. The thickness of the band is varied to Iimit the strain in the conductor to prevent degradation of its current-carrying capacity. The band is fastened at each end on each layer. This will enable tension to be developed within the band. The former thickness is sized to achieve the required prestress during cooldown to minimize conductor motion. When the coil is energized, the compression is relieved as the magnetic force in the conductor increases. This results in a relatively constant tension in the banding until the magnetic forces exceed the preload forces. This concept was used successfully in the 300-kJ coil [2 ]. The prestress in both the band and former are dependent upon the physical and mechanical properties of the two subassemblies. The technique assumes an infinitely rigid conductor. The concept also assumes that each layer acts independently. The former thickness can be structurally graded for each layer; however, for the conceptual design, a constant thickness was chosen for every layer. The materials selected for the structural component include G-10 Cl for the axial structure and spacers, KEVLAR for the former and 304L [3 ] stainless steel for the banding. Using a maximum bending stress of 38.6 MPa (40% of 96.5 MPa), assuming an axial alignment of fibers, the tooth thickness was calculated at 1.96 cm. The shear stress was calculated to be 5.8 MPa, which is 12% of the allowable value of 48.2MPa. The radial strain was limited to 0.1% for Nb 3 Sn superconductors. This resulted in band hoop stress of 201 MPa. The variation of banding thickness with radiuswas calculated and the results are shown in Fig. 2. Using these banding thicknesses and a former thickness of 1.07 cm, the banding prestress due to cooldown was calculated and the results are shown in Fig. 2. The axial structures are bonded to the formers using polyurethane adhesion. Polyurethane adhesives are the most appropriate for use at cryogenic temperature This is because they share the least embrittlement at very low temperatures compared to other classes of adhesives. They can be cured at room temperatures and do not need much pressure during eure. e]. Thennohydraulie Analysis and Design The forced-cooled design is very similar to the pool-cooled 300-kJ coil design [2 ] except that there are no bypass channels around the conductor and the conductor is cooled by crossflowing helium. Crossflow was selected instead of parallel flow for this coil on the following basis: 1. A sheathed cable could not be reliably bent around the small radii. 2. Headering of the coil would be difficult because of its compactness. 3. In parallel flow the pressure drop is inversely proportional to temperature, so a minimum additional flow over the basic stabilization value is required to maintain adequate flow to a normalized, elevated temperature section. 4. Heat transfer coefficients are high for small strands of conductor in crossflow. Coneeptual Design of a Snperconductinc Obmic-Heatinc CoU 109 l75r-------------------------------------------------------, .500 ~........ ..,e vi ...z ~ ( l25 I I I I D _;i I lDD u ,_ o. 75 :;: "'l!: 0 BANDING I ~;rHICKNESS I \ fVOD \ ,< I I .350 ii I .275:;; \.I I 0.25 I ~c- ..-c'\_BANDING\ I 0.00 I :i 0.50 "" / I 0 .425 \ PRE-STRESS 25 50 I ~ .,.~ "" .200 sJo COIL RADIUS. cm Fig. 2. Variation of banding thickness and stress with radius. 5. The helium expansion effect occurring during a normalization has a minimal effect on the ftow field in crossftow cooling. The Ieads from the coil to room temperature are conduction heat transfer dominated and cooled by supercritical helium ftow which is controlled by a critical orifice at room temperature. The high ftow rate of the 50-kA Iead precludes any significant rise in pressure in the expansion of the coolant owing to absorption of energy during a quench. A pressure of 6 atm and a temperature of 4.2 K were selected as the Operating thermodynamic state of the coolant, based on adequate thermodynamic-ftuid dynamic stability of the coolant channels and adequate dewar strength. A supercritical ftow circuit is shown in Fig. 3. The heliumpump and downstream heat exchangers are located in a dewar filled with helium by the refrigerator. The refrigerator capacity is matched tothedewar Ioad by the increase in heat leak as pool depth is increased. By applying this method of forced circulation, the refrigerator is decoupled from the coil, coil dewar, and Iead Ioad. Figure 3 also shows a cold Pressurlzer Une Regulator ,--------------0<1- From Relrigerator Cold Return to Relrigeralor J-T Vahe Dewar Warm Return Fram Leads Fig. 3. Forced-cooled coil ftow schematic. S. K. Singh et al. 110 Return ~ ~ Coil ( ( V ~ I Room Te mperature Critical Nozzle _,_....,-150gph ~alm b.. u::J ~ Dewar .. Sh 1elds Radlaiion Return Leads 4.2K ~ Lead Casing Fig. 4. Forced-cooled Iead and coil flow schematic. pressurized line used to make up the Iead ftow and to control the pressure in the forced-cooled circuit. The conductors are transverse to the axial ftow of coolant in the coil. Flow enters the coil on the centerline and ftows to a distribution header on the opposite end as shown in Fig. 4. The ftow proceeds to the opposite end where an insulating shroud returns the ftow over the coil surface eliminating any possibility of heat leak to the conductor from a stratified dewar. The ftow then turns to ftow upward along the dewar walls to carry away any heat leak from ambient and returns to the pump. This system prevents dewar stratification and the recirculation of any temperature spikes introduced into the coolant by pulsed operation. Stability Analysis In the initial design phase it was determined that for the conductors and geometries that are practical for this coil, a forced convection coefficient of 0.07 W I cm 2 - K is appropriate. Because a strand-to-strand insulation has not been identified for Nb 3 Sn, a representative value of k/ 8 = 1.0 W / cm 2 - K was used in conjunction with the critical temperature for I/ lc = 0.50 to determine that a heat transfer coefficient of 0.2 W /cm 2 -K is required. An insulation having a k/ 8 of unity could be a polyamide, polyethylene, or epoxy of 0.01-mm thickness or glass 0.1 mm thick. The peak Iosses in the forced-cooled design result in a surface heat ftux of 0.058 W /cm 2 or a local peak temperature of the conductor of 6.0 K. This is weil below the current-sharing temperature of 6.9 K. The average heat generation per cycle will not exceed 0.83 J/cm 3 of conductor or 0.577 J/cm 3 of helium in the conductor. lf the heat generated in a full bipolar swing is stored locally in static coolant the temperature will rise to 5.3 K. In a 1-s cycle the helium within the conductor volume will be replenished 50% from the coil channels so that satisfactory performance is assured. The residence time of helium in the coil is approximately 26 s which will not produce an excessive temperature rise in the shortest bipolar cycle of 11 s. Depending on the exact space-time averaging of the ac losses, the ftow velocity could be increased by 10-15% for safety margin, but the need for additional margin is not demonstrated by this analysis. The forced-cooled conductor stability is described in terms of a steady and transient criteria. Whereas the steady-state minimum propagating zone criteria for pool cooling is satisfied by a simple choice of conservative surface heat ftux, the forced-cooled conductor average surface heat ftux is lower, leading to a design minimum propagating zone that specifies the cooling system pumping power. Conceptual Design of a Superconducting Ohmic-Heating Coil 111 The transient stability of the conductor is described in terms of recovery from an instantaneous, adiabatic, 20-K conductor temperature rise with the surrounding helium at 4.2 K. The half-turn criterion is of sufficient length that axial conduction can be neglected. Cooldown is then given by simultaneous solution of (1) and (2): (T-t-q"'Acu/Up)/(To-t-q"'Acu/Up)=exp[- Up (-r--r0 )] CvPAc (1) This simplistic method of solution is adequate because the temperature rise of the helium is less than 2 K. The temperature of the conductor is found by marehing forward in small time steps and is conservative because the conductor-ftuid temperature difference is written in terms of the fluid temperature at the end of the time step, not the average as is customary. The transient heat transfer coefficient, allowing for surface renewal by turbulent eddies during the thermal boundary layer development process, is given as 1/2 k ] h = 2 [ pcP (3) 7T(T- To which decays to the steady value given by jh = 0.91 Re- 0 . 51 1/1, Re:::; 54.6 (4) jh = 0.61 Re- 0 . 4 1/1, Re 54.6 (5) h . - lh- CpGo (Pr 2:: )2/3 (6) "'= 0.91 (7) The overall U is determined as a series impedance of the Kapitza resistance (taken here as 2.5 cm 2 -K/W), insulation resistance (8/k = 1.0 cm 2 -K/W) and convection impedance (liters/hr). Application of the simple series impedance formulation can be verified by the heat transmissiontime constants for the conductor, insulation, and helium which are approximately 1,_"s, 0.1 ms, and 0.2 s, respectively. Solving (1) and (2) yields a recoverytime of 2.4 ms and the information that the required steady-state heat transfer coefficient is much less than that required by the minimum propagating zone (MPZ) value. In this case the transient stability criterion does not set the minimum ftow rate. The forced convection cooled coil is designed so that the Stekly energy balance criterion is exceeded in the conductor volume swept by the cooling channels and the steady-state minimum propagating zone for a crossftow cooled conductor given by (8) exceeds the width of the conductor support pad: kA) (:::; 2 ( Up 112 tanh - 1 ( 1 ) ------.,. 2- - - - - - - - [pef Acu/ Up(Tc - t)]- 1 (8) For the values in Table III the minimum propagating zone is 1. 79 cm, which is larger than the repeat length of the structure. From (5), the mass velocity required to obtain the space-averaged heat transfer coefficient of 0.032 W/cm 2 -K is 0.782 g/cm 2 -s. The total ftow area of the coil is S. K. SiDP et t1L 112 Table m. Parameten Used to Detennine the Minimum Propagating Zone Value Parameter U =(~X 0.003 +~X 0.09),* W/cm 2 -K Aeu,cmz A."cm 2 3 p 2 PeJ, W/cm Pe at 9 T, 0-cm Tc-t,K keuat9T, W/cm-K 0.0320 2.775 3.69 144.9 24.59 7.575 X 10-8 6.4 1.3 *Average over the swept conductor volume and still (supported) volume. 2.835 x 103 cm 2 foratotal ftow rate of 2.217 kg/s. The pressure drop is given by the Burke-Plummer relation, equation (9), for a Reynolds number of 8000: äpi = L 'f 1- e 3.5 _!_(G~) e3 Dp 2 p 9) ( The expansion and contraction Iosses per meter are äPi L [(1 - e )2 + (1.0/ e na~/2p L X (10) Fora coillength of 1.57 m the total pressure drop is 0.947 kPA. Pump work is given by the relation Wp = (rh/ p )M (11) '71th For a pump thermal efficiency of 50% the cold refrigeration Ioad due to pumping is 29.9W. The quench analysis of a coil is solely concerned with survivability of the coil and dewar. A conventional quench of the coil with the protection circuit and dump resistor Operation results in 0.03 J/cm 3 heat addition to the conductor. This is less than a conventional bipolar swing. CONCLUSION The design has demonstrated the feasibility of the ohmic-heating coil. The thermal analysis indicates that the cooling of the coil is adequate for the coil to remain superconducting during a full bipolar swing from +9 to -9 T. The structural analysis shows that enough structure is provided to withstand the Lorentz forces. In brief, the overall coil design meets or exceeds all the performance requirements. NOTATION Ac = area of conductor Aeu = area of copper CP = specific heat at constant pressure Cv = specific heat at constant volume Dw = diameter wire Conceptual Design of a Superconducting Ohmic-Heating Coil 113 G = mass velocity G 0 = superficial mass velocity h = heat transfer coefficient ih = Colburnj-factor J = current density . k = thermal conductivity L = length . M' = mass helium per unit length of conductor m = mass flow rate p = perimeter, pressure Pr = Prandtl nurober q"' = heat generated per unit volume Re = Reynolds nurober t, T = temperature of coolant, metal U = unit thermal surface conductance wp = pump work Greek Symbols 8 = insulation thickness e = void fraction = thermal efficiency = density p, = resistivity T =time ( = minimum propagating zone 11th p Subscripts c = critical o = outlet, outside s = surface REFERENCES L M. A Janocko, IEEE Trans. Magn. Mag-15(1):794 (1979). 2. S. K. Singh, J. H. Murphy, M. A. Janocko, H. E. Haller, H. Riemersma, T. L. Vota, D. C. Litz, R. Gromada, P. W. Eckels, Z. N. Sanjana, and F. N. Domeisen, "Prototype Tokamak Ohmic-Heating 20 MJ Superconducting Coil Study," Part I I -Technical Report, Contract No. L-48-8407C-1 (April 1978). 3. E. Mullan, D. W. Deis, P. W. Eckels, H. E. Haller, 111, M. A Janocko, S. A Karpathy, D. C. Litz, C. J. Mole, P. Reichner, Z. N. Sanjana, and M. S. Walker, "Design and Fabrication of 300 kJ-Superconducting Energy Storage Coil," E. M. 5077, Subcontract No. XN4-32767-3 (July 1977). 4. P. Thullen, J. D. G. Lindsay, D. M. Weldon and H. F. Vogel, "Superconducting Ohmic-Heating Coil Simulation," presented at Applied Superconductivity Conference, Pittsburgh, Pennsylvania, 1978. B-7 SHAPE OPTIMIZADON STUDY FOR A THREE-TUNNEL SUPERCONDUCfiVE ENERGY STORAGE MAGNET M. N. EI-Derini University of Petroleum and Minerals Dhahran, Saudi Arabia and R. W.Boom University of Wisconsin Madison, Wisconsin INTRODUCfiON lt has been shown CJ with the aid of the virial theorem that, regardless of configuration, an energy storage magnet would be prohibitively expensive if all structural supports were built with known fabricated materials. Accordingly, it has been proposed that the magnet be buried in tunnels in bedrock, a less expensive structure. For an underground solenoid it is relatively easy to carry radial pressure out to the bedrock. Struts connect the conductor to footpads on a rock wall so that radial forces are transmitted directly to the rock. Axial forces present greater difficulties because in a straight solenoid they are parallel to the walls. The compression Ioad produces a severe s}lear stress in the rock if this Ioad is carried directly to the wall. One would need a virial theorem mass if this Ioad is carried intemally in the fabricated structure. The concept of a segmented solenoid was introduced to help alleviate this problem. Cold structures may be used to carry axial Ioads in the upper segment of an hourglass segmented solenoid to the bottom of the segment. Similarly, cold structures carry forces to the ceiling of the lower segment. Axial forces are carried a short distance and transferred to a rock ftoor or ceiling through insulated struts instead of transmitting them from one end of the magnet to another by an expensive cold structure. There is still a substantial structure required to carry the forces along the conductor before they are transferred to the tunnelend or across the midplane of the centrat section. The mass of the structure can be reduced still further by increasing the number of segments. Unfortunately there must be a substantial rock wall separating each segment from its neighbors, otherwise there would be a danger of 114 Shape Optimization Study for a Supercondudive Energy Storage Mapet 115 Fig. 1. C-shaped solenoid of radii R 1 , R 2 and lengths h 1 and h 2 • collapse in the tunnel structure. Also, as the number of tunnels and the number of separate dewars increase, the size of a machine increases for a given energy. Constant-field magnets reduce the axial forces and the shear stresses on the walls. Thus, constant field segmented magnets will have both the advantage of constant field magnets and segmented magnets. An hourglass constant-field segmented solenoid is better than a straight constant-field segmented solenoid. Hourglass solenoids for energy storage use reduce the maximum stress of straight segmented solenoids from 6000 to 3200 psi for a 10,000-MW-hr magnet. A broad study for a three-tunnel solenoid has been performed to ascertain the best shape of a segmented solenoid which minimizes the axial forces and shear stresses. A new cross-sectional configuration in the form of a C, as shown in Fig. 1, seems tobe optimum. The length and radius of the central section are h 1 and Rh while the length and radius of the horizontal end sections are h 2 and R 2 , respectively. The distribution of the turns is not uniform but is distributed in a manner to obtain constant field along the inner surface of the solenoid. The ampere-turn per unit length which is proportional to the distribution of turns is shown in Figs. 2 and 3 for the vertical and horizontal sections. 3000.------,------,-------.-----. .. ·;;; CL CL Ul Ul 1&.1 Ul Ul 1&.1 a:: a:: I- I- U) U) -1000 -3ooo~----L---~L---~L---Jo 0.0 2.0 4.0 6.0 7.5 AXIAL DISTANCE , m Fig. 2. Ampere-turns per unit length, normal and shear stresses in the central section of C-shaped solenoid. 4 - 200 <>&e 10 12 74 75.5 HORIZONTAL DISTANCE , m Fig. 3. Ampere-turns per unit length, normal and shear Stresses in the end sections of the C-shaped solenoid. 116 M. N. EI-Dedlll _. R. W. Boom Fig. 4. A general three-segmented solenoid. The hourglass is for 8 = 90" and the C-shape is for 8 = o•. PROCEDURE The current distribution per unit length, J(z/ 1), is represented by Legendre polynomials as J(z/ I) = aopo(z/ I)+ azpz(z/ I) + a4p4(zj I) (1) where a 0 , a 2 , and a4 are evaluated using the Ieast-squares method to obtain a constant magnetic field along the inner surface of the solenoid. This procedure has been described in previous publications 4 ]. In addition to developing an optimum configuration for the constant field solenoid, a broad study of the three-tunnel solenoid (see Fig. 4) has been performed. In this study the hourglass and the C-shaped magnet are obtained when fJ = 90° and fJ = 0°, respectively. Results are presented for different fJ's, the same radii R 1 and R 2 , the same length of the centrat tunnel h 1 and the same maximum magnetic field, BM = 5.5 T. The length of the end tunnel, h 2 , is adjusted so that the stored energy, E, is the same in all solenoids. The axial force quality factor, the maximum shear and normal stresses vs. the angle, fJ, and the length h 2 are plotted in Figs. 5, 6, and 7 for e· Fig. 5. Axialforce quality factor for dillerent shapes from two dillerent views. Sbape Optimization Study for a Supercondudive Energy Storage Magnet 117 Fig. 6. Maximum shear stresses from two different views. constant energy with R 1 = R 2 = 70 m, h 1 = 7.5 m, and z 2 = 14m. lt may be shown that more length of the end tunnel is required for constant energy conditions if the angle 8 decreases from 90° to 0°. This means that more magnet area and more conductor are required for the C-shaped solenoid than for the hourglass solenoid. However, the C-shaped solenoid has smaller axial forces and smaller shear stress in the walls than the hourglass solenoid. The shear stress of the C-shaped solenoid is 22% less than that for a constant field hourglass solenoid. Another important point is that the structure of the C-shaped solenoid is a small fraction of the virial theorem mass, Oe. The latter factor is a compelling reason for selecting a constant-field C-shaped solenoid similar to that shown in Fig. 8. The quality factors of the C-shaped solenoids are calculated for different aspect ratios, ß. The radial, vector forces and mass factors ( 0 1,, 0 1", and Qc) increase for higher ß. The axial force factor, 0 1•• and the maximum shear stress on the other band, decrease for higher ß. The maximum normal stress has a local minimum between Fig. 7. Maximum normal stresses from two different views. 118 M. N. El-Derbli IUld R. W. Boolll Fig. 8. Mass quality factor for different shapes from two different views. aspect ratios of 0.19 and 0.30. The maximum shear stress occurs on the horizontal tunnels, while the maximum normal stress occurs on the vertical central tunnel. The dimensions of the tunnels and their relative positions affect the maximumnormal and shear stresses. To decrease the shear stresses, the outer tunnels should be close to the central tunnel and take into consideration the minimum distance between tunnels to prevent collapsing. CONCLUSIONS To make a large energy storage unit economical, the compressive forces must be reduced. lt is important to reduce the axial forces or the fractional virial theorem mass, Oe. This is achieved by using a constant-field C-shaped magnet as shown in Figs. 5 and 8 because the C-shaped magnet has the minimum 0 1• and a low Oe. lt is also important to reduce the shear stress to increase the design feasibility of the tunnels and reduce their cost. The C-shaped magnet has a minimum shear stress as shown in Fig. 6. The conclusion derived from these results is that the constant-field C-shaped magnet is a good solution for the shear stress and axial force problems. In addition, the constant field C-shaped magnet has the advantages of the constant magnetic field solenoid, i.e., it utilizes the superconductor more effectively and does not have the end field problem. A final important advantage of the constant-field C-shaped magnet is that tunnel excavation is relatively simpler. As noted earlier, the principal disadvantage of the C-shaped magnet is that it requires more conductor and larger surface area. ACKNOWLEDGMENTS This work has been supported by the National Science Foundation, Department of Energy, and the Wisconsin Electric Utilities Foundation. NOTATION a0 , a 2 , and a4 = Legendre polynomial coeflicients = maximum magnetic field E = stored energy h 1 = height of the central section of the C-sbaped solenoid h 2 = length of the horizontal end sections of the C-sbaped solenoid BM Shape Optbnizadon Stndy for a Snperc:ondnctive Energy Storage Magnet 119 J = current per unit length I = length of the magnet n = number of turns Oe= fractional virial theorem mass Q1, = radial force factor Q1• = total vector force factor Q1z = axial force factor R 1 = radius of the central section of the C-shaped solenoid R 2 = inner radius of the horizontal end sections of the C-shaped solenoid Z = axial distance from the median plane ß = aspect ratio = Iengtb/average diameter fJ = angle of the end sections on the horizontal plane REFERENCES 1. R. W. Boom, "Wisconsin Superconductive Energy Storage Project, Feasibility Study Report," Vol. I, University of Wisconsin,.Madison, Wisconsin (July 1974). 2. M. N. EI-Derini, R. W. Boom, and M. A. Hila!, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 88. 3. R. L. Willig and R. W. Moses, Jr., IEEE Trans. Magn. Mag-13(5):1122 (1977). 4. M. N. EI-Derini, Doctoral Dissertation, University of Wisconsin, Madison, Wisconsin (1978). B--8 20-kA POWER SUPPLY FOR LARGE SUPERCONDUCDVE COILS M. Masuda, T. Shintomi, and K. Asaji National Laboratory for High Energy Physics Ibaraki, Japan INTRODUCfiON There are no comQlercial power supplies designed exclusively for use with superconductors. Instead, conventional power supplies are slightly modified, taking into consideration the Ioad impedance of the superconductors. The recent trend has been that superconducting magnets carry large currents; for example, 20 kA in fusion reactors and more than 100 kA in energy storage coils. lt is doubtful whether conventional power supplies will be able to handle such large-current magnets. The purpose of the power supply for superconductors is twofold. One is for exciting superconducting magnets and the other is for supplying a current to short-length test samples of superconductors. Some typical component configurations of power supplies include: (1) a large-current transformer, a rectifier, and an array of transistors for current stabilization; (2) a large-current transformer, rectifier diodes, and thyristors connected to the primary winding of the transform er for current stabilization; (3) a large-current transformer and thyristors connected to the secondary winding of the transformer for current stabilization (the most conventional type); and (4) a large current transformer with a mechanical voltage adjuster connected to the primary winding of the transform er. Other variations can be considered including the motor generator. All of the conventional power supplies have problems when considered for use with large-current superconducting coils. The most serious problern is the use of a large-current transformer for impedance conversion. The secondary winding of a transformer carrying more than 20 kA is no Ionger fabricated of wires but of bus bars. The wide and thick conductors make its winding too difficult to be economical. Thus, the power supply for more than 20-kA output should be based on other design criteria. A new principle is introduced in the proposed power supply. Even so, a small current transformer is still employed to generate the dc source. The direct current is converted to ac by thyristor switching. The idea of dc conversion is not new. lt is commercially available as a so-called switching regulator, which is now used in place of a transistor-dropper method. The method introduced here uses switching but is quite different from the switching regulators. 120 121 20-kA Power Sapply for Luge Sapereoadaetive Coils I. 2~ 2~ L~ L~ Fig. 1. Elementary circuit for dc-dc conversion. BRIDGE 62 ~ ++ + CAPACITOR {~ 0 i~ l BRIDGE 63 PRINCIPLE OF OPERATION The basic principle of switching involved here is an ac-dc-ac-dc conversion. The commercial ac line is converted to a large direct current through successive processes of rectification, inversion, and rectification. The significant part of the process is the dc-ac-dc conversion through specially designed thyristor units [ 1]. The conventional dc-ac inverter using thyristors is commercially available and is designated as a voltage-controlled inverter. The waveform of the output voltage is rectangular. On the other hand, the current-controlled inverter is a device commonly used to control the speed of induction motors. The device introduced hereisdifferent from both types. Since commutation is performed by capacitors charged by the constant current, this is a current-controlled inverter. The significant characteristic of this circuit is current amplification, which makes it effective as a high-current dc power supply. The principle of the current amplification is given below but a detailed analysis of the circuits is beyond the scope of the present paper. In Fig. 1, 10 is the constant input dc current. The current ftows in the thyristor bridge, B2 as in a conventional, three-phase Graetz bridge. The ac voltage is generated on terminals of capacitors which cause thyristor commutation. The waveforms of the voltages generated for delta and star capacitor connections are shown in Figs. 2a and 2b. The ac voltage is rectified to supply the dc current to the Ioad. The firing of bridge B3 is in the same sequence as in B2 but at different angles. The output voltage is controlled by the firing angle of B3. By rough estimation, the peak voltage at the capacitor, Vc, is given as Vc = kloi/C (1) where f is the inverter frequency in Hertz, C is the capacitance in farads, 10 is the input current in amperes, and k is a constant. At no Ioad, Vc is controlled by 10 • At full Ioad, the input and the Ioad currents ftow together into the capacitor and Vc increases. When the states for no Ioad and full Ioad are defined as 0 and 100, respectively, and the ac voltages are given by Vc(O) Vc(100) = klo/f(O)C (2) = kiJf(100)C (3) where h is full-load current. It is assumed that IL » / 0 • From (2) and (3), the following equation is obtained: IJio = [/(100)//(0)][Vc(lOO)/Vc(O)] (4) When 10 is held constant, Vc(O) is generated essentially from the forward voltage drop of the thyristors and is approximately equal to the input dc voltage. The voltage, 121 M. Masuda, T. Shintomi, and K. Asaji VcJ ~ ""-'-/ (a) (b) Fig. 2. (a) Calculated waveform of voltage generated on capacitors which are connected in a star configuration. (b) Calculated waveform of voltage generated on capacitors which are connected in a delta configuration. Vc, increases with load current to a maximum value of Vc(100). The required output power, P, is given as (5) where K varies slowly with h. The input voltage changes with the load while the input current is kept constant. Equation (4) shows the current amplification. For example, if the load current is 20 kA and the output voltage is 20 V, then the outputpower of 400 kW could be supplied from a 200 A and 2 kV input. The current amplification, h/ Io, is 100 and this is achieved by a voltage ratio Vc(100)/ Vc(O) of 20 and a frequency ratio /(100)//(0) of 5. The frequency ratio of five implies f(O) = 100Hz and /(100) = 500Hz, for example. In the present inverter, as shown in Fig. 2, the thyristors employed do not require fast switching characteristics because the commutation is performed by the slowly charged voltages on the capacitors. The maximum frequency possible with conventional power thyristors is about 1 kHz. MODEL POWER SUPPLY A 1000-A modelpower supply was constructed and excited with a 100-kJ superconducting magnet [2 ]. The approximate specifications of the model unit are given in Table I. The schematic diagram of the control unit is shown in Fig. 3. The master oscillator supplies a high-frequency clock whose frequency is divided by frequency dividers. Their outputs are fed to distribution circuits com~sed of shift registers and power amplifiers. The thyristors of bridge B2 are fired sequentially with phase differences of 60°. Bridge B3 is fired in the same mode as B2 but with a phase difference from B2 that is determined by counting a preset number of cycles of the 123 20-kA Power Supply for Large Superconductive Coils Table I. Specifications of Test Arrangement Input dc power supply Thyristor of bridge 2 and 3 Commutation capacitor/arm Connection of capacitance Dummy Ioad magnet Superconductive Ioad coil 100V,50A 1200 V, 1000 A 200 V, 500 f.LF X 5 Star 5 mH, 10mn 1000 A, 0.2 H, Nb-Ti master oscillator frequency. The phase difference between B2 and B3 is independent of the frequency of the master oscillator which enables the Operation of the inverter to be very flexible. As shown in (4), to obtain the highest current amplification, the lowest possible valtage Vc(O) must be employed. In that case, thyristors tend to miss the commutations because of low anode voltages. The low valtage is compensated for by decreasing the master oscillator frequency with the variable frequency controller VFC. There are two methods to control the output current. One is with the phase angle and the other is with the frequency. Figure 4 shows the waveform of the ac valtage and the current in the capacitor in the delta connection. Figure 5 shows the output current vs. the phase angles. In the present work, the electromagnet constructed of a normal conductor was used as a preliminary Ioad. Aftertuning the model unit, the 100-kJ superconductive magnet was connected and excited to 1000 A. The specifications for the test are shown in Table I. The efficiency of the model unit is about 7 5%. Since the lass by the thyristor forwardvaltage drop is dominant, the efficiency can be improved more than 90% by improved design of the power supply, which has a sufficiently higher outputvaltage than the forward voltage. DATA I-~- INPUT(f) DATA INPUT(8) Fig. 3. Schematic diagram of the control circuit. SCR'S TRIGGER CIRCUIT 124 M. Masuda, T. Shintomi, and K. Asaji Fig. 4. Measured waveform of generated voltage and current on andin the capacitors. CONCEPTUAL DESIGN OF A 20-kA POWER SUPPLY The conceptual design of a 20-kA power supply has been carried out. No special parts have been used and no technical difficulties were encountered. The connection diagram is shown in Fig. 6 and the specifications are given in Table II. The advantages of the proposed power supply are as follows: 1. 2. 3. 4. lt is easy to generate a large dc current. The system does not need a large current transformer. Only a small section of busbar is needed. By employing frequency changes, the thyristors can operate without commutation failures because of sufficient anode voltage. 5. The control unit may be digital and thus noise resistant. 6. The unit is essentially a constant-current source, providing safety from overcurrent caused by commutation failures and Ioad short circuits. INPUT CURRENT 20 A 200 10 A ... ~ !5 u 100 SA ~ 30 90 60 FI RING />NGLE , 0 Fig. 5. Load current vs. the firing angle of the thyristors in the bridges. 20-kA Power Supply for Large Superconductive Coils l V. AC INPUT 3LOAD I ill BRIDGE 81 BRIDGE 82 CAPACITOR 125 BRIDGE 83 Fig. 6. Eleroentary circuit of the 20-kA power supply. 7. 8. The operation frequency can be high even with conventional thyristors providing means for reducing ripples present in the output current even with small filters. The harmonic current generated from the thyristor instruments has frequencies of (n ± 1)/ where n is the pulse number of the rectifier; therefore, there will be little pollution of the commercial ac line. There are two conventional types of power supply for superconductors on the market. One is the large current transformer with gate-controlled thyristor bridges for current regulation. This type is valuable for large power but not for the large currents that are required for a superconductive energy storage unit for diurnal energy storage on a utility network nor for the huge coils of fusion reactors. The multiphase power supply has a ripple current with characteristic fundamental frequencies of nf, 300Hz, or 600Hz in a 50-Hz utility system where n is the pulse number of a thyristor bridge and 6 or 12 in general. However, a slight unbalance between voltage and phase of the three phases or an unbalance between firing circuits will give noncharacteristic ripple frequencies such as 100, 150, 200Hz, etc. in the output current. For the characteristic frequencies, the maximum ripple, for example at 300Hz, is about 5% when thyristors are fired with the maximum angle. Decreasing the ripple to 0.1%, requires an attenuation in the filter of around 34 dB. When the attenuation rate of 6 dB/oct of the conventional ripple filter composed of inductors and capacitors are employed for ripple reduction, the resonance frequency Table II. Appro:ximate Specifications for 20-kA Power Supply Input transforroer Thyristor in bridge 1 and 2 Cororoutation capacitor Thyristor in bridge 3 Output Priroary Secondary Power rating Voltage Current Nurober Voltage Capacity Nurober Voltage Current Nurober Voltage Current 6.6kV 660V 400kVA 2.0kV 200A 12 800 V dc 330 V ac 14roF 3 2.0kV 2500 A (3 phase) 18 15V 20kA 126 M. Masuda, T. Sbintomi, and K. Asaji of the filter must be lower than 10Hz. For the worst case, that is, for noncharacteristic frequencies, the size of the filter is not economically acceptable. The other power supply, which also has a large-current transformer, the rectifier with parallel-connected transistors for current regulation and for ripple reduction is also not acceptable because employment of a number of transistors makes it less reliable. The proposed power supply can use an operational frequency higher than 500 Hz without expensive high-power thyristors specifically designed for fast switching, and a small filter can adequately reduce the ripple current. HEFERENCES 1. H. A. Peterson, M. A. Hilal, W. C. Young, and R. W. Boom, in Energy Storage, Compression, and Switching, Plenum Press, New York (1974), p. 309. 2. M. Masuda, T. Shintomi, S. Matsumoto, H. Sato, and A. Kabe, in Proc. 6th Intern. Conference on Magnet Techno/ogy, ALFA, Bratislava, Czechoslovakia (1978), p. 254. C-1 SUPERCONDUCTING GENERATOR DESIGN FOR AIRBORNE APPLICATIONS* 8. 8. Gamble and T. A. Keim General Electric Company Corporate Research and Development Schenectady, New York PROGRAM OBJECfiVE During the past eight years, the United States Air Force has actively pursued the development of lightweight, superconducting generators in the multimegawatt power range for airborne applications [ 1 ' 2 ]. This development work coupled with that of university and industrial laboratories for utility applications has pointed to the technical feasibility of these machines. Extrapolation of this technology to higher power densities requires the combination of advanced component concepts and materials in a single machine. The objective of the design presented here is to consolidate advanced concepts and advanced materials developed for superconducting generators to attain specific design goals and to advance the state-of-the-art for superconducting generators for airborne applications. DESIGN GOALS AND CONSTRAINTS The airborne high-power system is depicted schematically in Fig. 1. The generator is to be designed to provide 20-MW of dc power at a voltage of 20-40 kV from the rectifier. A nominal speed of 6000 rpm was selected from the limited speed range of the proposed drive turbine. The inertia is required tobe 13.6 kg-m (10 slug-fe) or Iess, and the specific weight goal is 0.045 kg/kW (0.1 lb/kW). These ambitious goals aretobe met in a design capable of 1-s field ramping and 1-s spin-up FUEL - Fig. 1. Airborne high-power system. * Work supported by Air Force Aero Propulsion Laboratory, Wright-Patterson Air Force Base, Ohio, under U. S. Air Force Contract No. F33615-76-C-2167. 127 1ll B. B. c..111e ud T. A. Keim Table I. Selected Design Cbaracteristics 20MW 4pole 6000rpm 200Hz 29.6 kV line-to-line (nns) Parameters Field current density, A/cm2 Field current, A Field inductance, H Armature current density (annulus), A/cm 2 Copper volume fraction (armature) Synchronaus reactance, pu Transient reactance, pu Subtransient reactance, pu Equivalent power factor Efficiency 15,000 867 0.69 1760 0.39 0.56 0.50 0.26 0.86 0.93 to design conditions. The generator is to be capable of both continuous operation with run times of up to 5 min and pulsed operation. OVERALL DESIGN These goals were used as design constraints in a machine sizing study to consider the most appropriate machine configuration for this application. The result of this study was the selection of a four-pole, conductively shielded machine with a direct water-cooled, involute end-turn armature. The generat characteristics of the selected design are presented in Table I. A cross-section of the overall design is presented in Fig. 2, and the corresponding critical dimensions are given in Table II. To achieve this compact, lightweight Table U. Selected Design Weight* lb kg Rotor Field-winding outer radius Shield outer radius Thickness Length Bearing to span 870 395 Stator Armature thickness Outer radius Shield inside radius Thickness Straight length Total 870 395 1740 790 * Dry weight, without auxiliaries. Dimensions in. m 9.0 10.5 0.75 21.0 32.0 0.229 0.267 0.019 0.533 0.813 1.0 12.0 19.7 0.75 6.0 0.025 0.305 0.500 0.019 0.152 Fig. 2. Design (cross section). TOROUE TUBE EXTENSION COOLANT HEADERS ENVIRON MENT AL SHIELO ELECTROMAGNETIC SH IELD = DRIVE FLANGE BE.ARING FILAMENT WRAP ~ .... Sl t f J Ii J· t 111 1 i{' 130 8. 8. Gamble and T. A. Keim design, the following significant design features have been included: 1. Field winding composed of four racetrack-shaped, epoxy-impregnated modules manufactured with multifilamentary Nb 3 Sn superconductor in a cabled configuration. 2. Nonconducting composite support structure shrunk between the bore tube and the torque tube. 3. Torque tube cantilevered from the undriven end of the machine. 4. Outer aluminum electromagnetic shield with a high natural frequency in the ovalizing mode. 5. High current density, direct water-cooled armature. 6. Conducting (aluminum) environmental shield to Iimit the fields external to the generator. 7. Integral water tank. ROTOR DESIGN The rotor design is most strongly influenced by the combined requirements of a 1-s start-up and a 0.1-lb/kW specific weight. The initial phase of this program included a sizing study which encompassed the selection of the superconductor, the principal rotor details, and the generator configuration. Superoonduc:ting Field Winding The field winding is the most critical rotor component. The field-winding configuration dictates the support structure design and much of the remaining rotor design. The field winding was designed on the basis that the superconductor must be constrained to prohibit wire-sliding motion in response to sudden field ramping and torsional spin-up. Therefore, a fully impregnated winding was selected. The potential Iimitation to the use of fully impregnated modules in fast-ramping applications is the temperature rise within the module, owing to heating of the transient conductor. Early in this program, it was decided to Iook to Nb 3 Sn with its high critical temperature as the means to achieve fast-ramping capability in a fully impregnated winding. The selected superconductor is shown in Fig. 3. This Nb 3 Sn cable is being developed under a United States Air Force contract at the Intermagnetics General Corporation. As discussed above, this conductor was selected primarily for its temperature capability. A completely fair comparison between the types of superconductors is difficult, but the generat magnitude of the difference can be indicated by a simple comparison. For the 6.8-T peak field and the 15,000-A/cm2 module current density of the selected field-winding configuration, Nb 3 Sn is capable of operating at about 8 K while NbTi would be capable of operating at temperatures less than 6 K. For an epoxy-impregnated field winding in a 4-K helium pool, this corresponds to a fourfold increase in enthalpy margin. The characteristics of the selected superconductor have been the subject of substantial investigation including both critical current optimization and loss measurements [3 ]. A cable is preferred for this application because it has a low-loss and high-current configuration. High current is desirable to Iimit the voltage required for rapid field ramping. Even with the superior performance of Nb 3 Sn, the large forces to be experienced by the field winding in this application dictate the selection of a winding design which Iimits local wire-sliding motion, frictional heating, and subsequent Superconducting Generator Design for Airborne Appliations 131 0.0008 in. BRONZE and Nb3Sn CORE Ta (7.2%) (44.4%) 0.039 in. /--... _....\. \ /~ \ (? ) '----- / /-...... \ MOLYBOENUM CENTAAL STRAND ( \( ....__/ "l / \ ) \ ......__...,- ) = 0.211 BRONZE AND Nb3Sn MODULE PACKING FACTOR OVERALL CURRENT DENSITY = 15,000 Alcm2 BRONZE AND Nb:JSn CORE CURRENT DENSITY = 7.11 x 104 Alcm2 8.4K (at 6.8 tesla) T0 9K (at 6 tesla) 0. 23 Cu FAACTION OF THE MODULE 864.6 amperes CABLE CURRENT = = = = Fig. 3. Preferred superconductor. quenching of the winding. It is the objective of the winding development portion of this program to combine the advanced Nb 3 Sn-cabled superconductor with the epoxy-impregnation techniques which have been successfuily applied to NbTi windings [4 ]. A racetrack shape with a reetangular cross-section was selected for a fieldwinding module configuration (Fig. 4). This configuration has been selected for its simplicity, which permits a more readily controilable manufacturing process than would be possible with a more complex shape. A module bore of 7.62 cm was selected to Iimit the conductor strain in bending to approximately 1%. The selected operating current density of 15,000 A/cm 2 , based on the crosssectional area of the winding module, results in the field distribution plotted in Fig. 4 for the midplane of the module at the start of the circular end-turn. The peak field seen by the superconductor is 6 .8 Ton the winding bore. This field quickly tapers towards zero near the module center. The peak field region is weil cooled, since it occurs on the module surface. The critical temperature distribution and the calculated temperature distribution foilowing a 1-s ramp to design current are also plotted in Fig. 4. As would be expected, the temperature distribution peaks weil away from the peak field location. Similar calculations indicate that a steady-state uniform heat rate up to 4 mW/cm 3 can be accommodated without exceeding the local critical temperatures. Aspart of the present program, verification of winding performance by testing a cylindrical winding module prior to manufacturing the racetrack modules is planned. The cross section and the bore of the test module are the same asthat of the racetrack modules. Testing will include rapid field ramping to verify transient performance. 131 B. B. G81Dble and T. A. Keim 13 12 11 "'· w cc ::> 10 9 1-- <{ cc w 8 ~ w 1-- 7 0.. K ~ 1.6 x 10""' w/ cm• K p cp = 6.06 x lO''T'·"Jicc • K @1 -4 1-- 7.62 cm(3 in. I ITB( 37.3cm (14.7 in.l ~"!tl (8.7 in. I 10 9 8 7 6 B(TI 5 4 6 3 5 2 4 0 L-~--~2---3~~4L_~5--~6---7L-~8--~9~ 0 X,cm 0 Fig. 4. Winding module field and temperature distributions following a 1-s ramp. Support System The function of the support system is to support the field winding in response to centrifugal and electromagnetic forces. The support system is depicted in Fig. 5. The racetrack-shaped modules are contained in a support structure constructed from G-11 CR epoxy-glass Iaminate. This grade specification was selected for this application because at 4.2 K it has a high modulus, 3.9 x 104 MPa (5.7 x 106 psi, in the warp direction) [ 5 ], and a high compressive strength, 724 MPa (1 05,000 psi, in the warp direction) and because it is nonconducting, it is lossless in an ac magnetic field. This structure is assembled around an Inconel-718 bore tube (23.7-cm OD, 19.13-cm ID). The function of the bore tube is to serve as an inner vacuum barrier and to provide radial stiffness for the assembly. The bore tube, the support structure, and the winding modules will be cooled to liquid nitrogen temperature and assembled inside the Inconel-718 torque tube. Inconel-718 is selected because it isahigh strength (uy = 1405.8 MPa, 203.9 ksi at 4 K), weldable, nonmagnetic material available in tubular forgings. The torque tube is to be prestressed to a tensile stress of approximately 60,000 psi to limit the relative sliding motion of the support structure with respect to the winding modules with spin-up and application of Ioad. An additional benefit of prestressing is to limit the magnitude of the cyclic stress seen by the torque tube. With the application of centrifugal and electromagnetic load, the torque tube experiences an overall hoop stress of approximately 620.5 MPa (90,000 psi) with a superimposed bending stress resulting in an outer fiber peak stress of 758 MPa (110,000 psi). Considering the entire cycle of spin-up, excitation, and shutdown, the torque tube will experience cyclic stress with a magnitude of 172 MPa (25,000 psi) on a base eJ. Superconducting Generator Design for Airborne Applications SUPPORT STRUCTURE 133 TORQUE TUBE FIELD WINDING MODULE Fig. 5. Field-winding and support system. stress of 386 MPa (85,000 psi) for a maximum of 72,000 cycles in 100 hr of operation. lt should be noted that there are no torque transmitting keys in the torque tube to cause stress concentrations. The torque is transmitted by friction between the support structure and the torque tube. A coefficient of friction of 0.045 is sufficient to carry the fault torque. The low-temperature region of the rotor is vacuum supported from the undriven shaft ftange by a single torque tube extension. The cantilevered design was selected to separate the axial thermal contraction and torsional deftections of the torque tube from those of the electromagnetic shield. In addition, this design reduces the heat input to the low-temperature region and greatly simplifies the assembly. The short length of the rotor allows this designtobe considered. The first rotor bending critical speed occurs at 1.5 times the speed of rotation. Electromagnetic Shielding The rotor shielding is designed as a two-component system. An ambient temperature electromagnetic shield excludes the majority of the nonsynchronaus fields from the inner colder members. In this application, an ac generator feeding a rectifier Ioad, the dominant source of the nonsynchronaus magnetic field, is expected to be the nonsinusoidal nature of the armature currents, resulting in a 1200-H frequency fieldas seen by the electromagnetic shield. A second actively cooled shield at approximately 100 K intercepts thermal radiation and provides additional electromagnetic shielding. The selected electromagnetic shield design is a 707 5-T6 aluminum, rolled-ring forging with a 53.6-cm OD, a 49.8-cm ID, and a 53.3-cm unsupported length. The shield thickness, 1.91 cm, presents 5.5 skin depths at 1200 H and was selected primarily from considerations of fault crushing forces and natural frequency in the ovalizing mode. The shield thickness results in a natural frequency of 1350 Hin the 134 B. B. Glllllble and T. A. Keim Table lU. Fault Conditions Associated witb Field Design Fault shear stress (4 pu) Fault peak bending stress Hoop stress in response to uniform fault compressive force Centrifugal hoop stress MPa psi 15.1 330.3 -62.2 78.4 2,189 47,907 -9,017 11,375 four-pole ovalizing mode, assuming fixed ends and neglecting electromagnetic stiffening effects. The fault conditions on the shield are summarized in Table 111. Shield heating is a particularly important consideration in this application. Prolonged single-phase faults cannot be tolerated, and it has been proposed to short-circuit all three phases of the armature in the event of a single-phase fault. A second area of particular concern is shield heating in response to nonsynchronaus armature currents associated with the rectifier Ioad. Shield heating computations are somewhat inexact, but it appears that in the absence of fittering or active shield cooling continuous operation may be permitted for 2 min. Analysis indicates this can be extended by blowing airdown the gap, by including cooling tubes on the armature bore, or by including a modest filter (mass 45 kg) in the system. Helium Flow Circuit The function of the helium ftow circuit is to remove ramping and other winding Iosses and to intercept heat conducted and radiated to the low-temperature region of the rotor. The winding modules are cooled by helium in the slots and channels in the support structure. A liquid-vapor interface occurs near the radially innermost winding module surfaces. Therefore, the majority of the winding is in contact with subcooled and supercritical liquid. In the high centrifugal field, large convective heat-transfer coefficients can be expected, and the thermal impedance to the transfer of heat to the liquid is dominantly in the impregnated winding module. The majority of the helium ftow circuit is located in the removable subassembly in the undriven end of the rotor. The ftow circuit components are depicted in Fig. 6. Liquid is fed into the rotor through the relative motion bayonet in the radial tubes feeding the winding region. This radial tube desigt! follows that developed for an earlier 20-MVA superconducting generator rotor [6 ]. The hydrostatic head of the liquid in the winding pool covering the ends of the radial supply tubes is balanced by that of the vapor column in these supply tubes to provide a ftow or Ievel control. The vapor generated at the winding pool surface at point 8 is compressed as it ftows radially to state point 10. The gas is heated as it ftows to state point 11 while cooling the torque-tube-extension heat exchanger. The ambient temperature gas ftows radially inward to point 12 before exiting out the helium transfer coupling. The compression from state points 9 to 10 is only partially offset by the expansion from points 11 to 12. The selected dimensions result in a pressure at the liquid-vapor interface of 0.42 atm and a temperature of 3.4 K. The winding-pool temperature distribution in this application is approximately isentropic in that the support structure and winding assembly have a low radial thermal conductance. The winding-pool temperature distribution varies from 3.4 K at the Iiquid-vapor interface to 4.4 Kat the outer surface. The steady-state mass flow is 0.6 g/s. Superoonduding Generator Design for Airborne Applicadons 135 ELECTROMAGNETIC SHIELD RETURN HELIUM LIQUID HELIUM ® Fig. 6. Flow circuit schematic. During transient operation, 1-s ramping Iosses (510 J}, sudden loading Iosses (30 J), and viscous Iosses (120 J) are introduced to the low-temperature region components. This serves to slightly pressurize the winding pool containing 4.5 Iiters of liquid. If the generator is operated in a continuous mode, this is relieved by an increase in ftow through the torque-tube-extension heat exchanger. If the rotor speed is reduced, the vapor evolved is vented out the exhaust tubes. These exhaust tubes are also intended to carry cooldown ftow. 1t should be noted that the torque-tube-extension heat exchanger is conical in shape to prevent axial heat transfer via secondary ftows in the cooling stream. The vapor-cooled, field-winding current Ieads slant radially inward for the same reason. ARMATURE DESIGN CONSIDERATIONS To date, the major thrust of the program has been in the areas of generator overall design and rotor detailed design, but the achievement of the goals of this program will also require significant advances in armature design. The most important armature design considerations are the requirements for 29.6-kV line-toline rms and the annulus current density of 1760 A/ cm 2 • The required voltage will be achieved in part by virtue of the short required life at voltage, 100 hr. The iron-free designwill utilize voltage grading to reduce the insulation requirement. The required current density will be achieved by constructing the armature conductor from hollow-copper which will be cooled by high-velocity, high-pressure, subcooled water. FUTURE WORK The achievement of the goals of this program requires the combination of advanced design concepts and advanced materials. The design presented here addresses these goals and will be developed as part of the present program through 136 B. B. GluOie Md T. A. Keim static testing of the rotor and initialarmatme component development. This includes the following: 1. Manufacturing a cylindrical module of a representative cross section to verify the pedormance of an epoxy-impregnated Nb 3 Sn cable winding for fast field ramping applications. 2. Individual pedormance testing of the winding modules before assembly into the rotor support system. 3. Static (nonrotating) testing of the rotor ftow circuit and the field winding while the rotor is suspended from the undriven end. The one torque-tube-extension heat exchanger design allows operation of the ftow circuit in this position. 4. Building armature bars and testing thermal pedormance and voltage capability. The selected design concepts will be verified with these tests. This verification will demoostrate that superconducting generator technology is extrapolatable to higher power densities than have been demonstrated to date. ACKNOWLEDGMENT The assistance of H. L. Soulhall of the Air Force Aero Propulsion Laboratory at Wright-Patterson Air Force Base, Ohio, is gratefully acknowledged. REFERENCES 1. H. L. Southall and C. E. Oberly, IEEE Trans. Magn. Mag-15:711 (1979). 2. R. D. Blaugher, J. H. Parker, and J. L. McCabria, IEEE Trans. Magn. Mag-13:755 (1977). 3. G. R. Wagner, S. S. Shen, R. E. Schwall, A. Petrovich, and M. S. Walker, IEEE Trans. Magn. Mag-15:228 (1979). 4. M. J. Jefteries and E. T. Laskaris, "Rotor Winding Development for a 20-MVA Superconductive AC Generator," paper presented at Conference on Technical Applications of Superconductivity, Alushta, USSR, 1975. 5. J. R. Benzinger, "Manufacturing capabilities of CR Grade Laminates," paper presented at NBS-DOE Workshop, Vail, Colorado, October 24-26, 1978. 6. E. T. Laskaris, IEEE Trans. Magn. Mag-13:759 (1977). C-2 SUPERCONDUCTING FJELD WINDING FOR A 10-MVA GENERATOR* K. A. Tepper, J. V. Minervini, and J. L. Smith, Jr. Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCTION A research program funded by the U. S. Department of Energy has been underway since 197 6 to develop advanced concepts for superconducting generators to be used in large central power stations. The background and scope of the program have been reported previously [ 1]. A prototype 10-MVA machine is being constructed and will be tested under fullload while connected to a power system. The first part of the present program has concentrated on the analysis of advanced concepts to solve key problems identified through earlier experiments and analyses e·3]. One major new feature selected for demonstration in the 10-MVA generator is a dual shielding system to protect the superconducting field winding from ac magnetic fields while still providing adequate damping. The generator Iayout is shown in Fig. 1. Both the outer wound damper shield and the inner solid conducting shield are positioned on the rotor and maintained at liquid helium temperature. The compact cryogenic structure for the shields minimizes the distance between the field winding and armature winding and thus maximizes magnetic coupling. FIELD-WINDING DESCRIPTION The support structure for the superconductors in the winding space is a multiplicity of yokes (see Figs. 2 and 3), each with a boss and a pin for support at the top and bottom. The side walls of the yokes are tapered, increasing in width toward the top, and the top and bottom walls are deeper than the side walls. The superconductors are wound into modules which are surrounded by the yokes. The modules are fitted into the rotor such that the yokes mesh (see Fig. 4), filling as much of the space between the inner and outer support cylinders with superconductor or structure as possible. The bosses at the top and bottom of each yoke contact the inner and outer support tubes. The yokes are supported against azimuthal and axial motion by pins inserted between the inner support tube and the yoke bottom, and the outer support tube and the yoke top. Once the yokes are filled with the superconductors, the yoke tops are welded onto the side walls. The weid is made to penetrate to the stress-relieving holes, which * Work supported by U. S. D.O.E. under Contract No. EX-76-A-01-2295 Task Order No. 11. 137 ROTOR SHAFT INNER SUPPORT TUBE OUT ER SUPPORT TUBE DAMPER Fig. I. Cross section of 10-MYA rotor. YOKE LHE RESERVOIR SUPERCONOUCTOR MODULE CENO TURNl WINDING SPACE .... ~ 1'1.> J r i!"" I ~ !"" ~ i p:: ?" ~ Superconducting Field Windlogfora 10-MVA Generator 139 .0 15 R. TYP. (8) PLC S. MODULE I 2 3 4 5 6 7 G 1.711 1.817 1.688 1.634 1.450 1.412 1. 268 H I 1.1 93 .259 1.193 .312 1.049 .320 1.04ll .293 .905 .273 .905 .254 .761 . 254 J K .743 .714 .743 .714 599 568 568 599 .484 .455 .4 24 .455 .279 .3 11 L .892 .985 .887 .911 .781 .805 .66 1 Fig. 2. Shop drawing of yokes. puts a radius at the end of the parting-line crack. The weid is made with the aid of a water-cooled chill to control thermal darnage to the superconductors. During a test weid, copper strips covered with epoxy were clamped to the yoke wall to simulate the superconductors. The weid was carried out successfully without any thermal degradation of the epoxy. The individual modules of superconductors are wound into radial layers with adjacent radiallayers separated by parting strips to form slip planes. The slipplanes Fig. 3. Photograph of yoke. K. A. Tepper, J. V. Minervini, and J. L. Smith, Jr. 140 / / / / / ----- - -- OUTER SUPPORT - -/ /~ I I r---?--~~~ I SUPERCONDU CTOR MODULES Fig. 4. View of straight section of winding space. are necessary in order to ensure that the epoxy superconductor interface shear stress is maintained below the failure stress of the bond. The two saddle coils have seven modules each, with the size of the modules varying with angle. The largest modules are at the junction of two saddle coils, and the smallest are closest to the pole faces. This shaping serves three purposes. First, it puts the maximum concentration of superconductors where they contribute the maximum to the dipole field; second, it puts the maximum concentration of superconductors where the azimuthal-directed self-forces are smallest; and finally this distribution minimizes the field concentration near the pole faces and in the end-turns. CONDUCfOR The conductor selected for the field winding of the 10-MVA generator is a monolithic multifilamentary composite conductor, 2.72 x 1.49 mm, with a copper matrix and niobium-titanium superconducting filaments. Two superconductors with these specifications are being considered. The copper-to-superconductor ratios are 1.65: 1 with 2100 filaments of 30 #Lm diameter and 2: 1 with 817 filaments of 46 #J-m. The 1.65 : 1 wire has a higher critical current and a lower hysteresis loss than the 2: 1 wire, which has a slightly higher thermal margin. Two test solenoids are being fabricated to evaluate the conductor performance under simulated operating conditions. Other important conductor and winding-module characteristics are listed in Superconducting Field Winding for a 10-MVA Generator 141 Table I. Field-Winding and Conductor Characteristics Conductor 1.65: 1 Dimensions, mm Copper: Nb-Ti ratio Filament diameter, 10-6 m Number of filaments Twist length, mm Operating current, A Conductor current density, A/m 2 Overall current density, A/m 2 Winding Ampere-turns Peak field, T Inductance, H Modules Dimensions, mm Number of modules Turns per module Maximum Minimum Total number of turns 2.72 X 1.49 1.65 30 2160 12.7 939 2.32 X 108 1.28 X 108 2:1 2.72 1.49 2 46 817 12.7 939 2.32 X 10 8 1.28 X 10 8 X 1.74 X 10 6 4.8 2.1 40.72 X 15.21 7 140 56 1456 Table I. This type of conductor was selected for the characteristics of low ac loss, ease of winding into the module structure, and mechanical and electrical stability. STRUCI'URAL ANALYSIS The major motivation for developing the yoke concept for the superconductor support system was to isolate the superconductors from the high shear stresses present in the inner and outer support tubes during the high-torque phase of an electrical fault. The stress and the deformation of the rotor were found by modeling the rotor as a circular cylinder carrying torque. The stress in the outer and inner supports are well within their shear strengths. In an early design the field-winding region between the inner and outer support was filled with alternating layers of epoxy-bonded double-pancake windings and epoxy fiberglass bands, a design very similar to the MIT 3-MVA rotor [4 ]. This design was unsatisfactory because the deformations of the center and outer supports are imposed on the superconductors in the winding space as an axial gradient of circumferential motion. All the superconductors are continuously epoxied together so they deform as a circular cylinder. The imposed deformation causes a shear stress at the epoxy interfaces between superconductors which exceeds the shear strength of epoxy. Note that the stresses in the epoxy are not because the composite winding carries a large percentage of the torque, but because the deformations of the inner and outer supports are imposed on the winding. The yoke-module configuration solves this problern since the circumferentially independent design of the structure isolates the superconductors from stresses developed during the high-torque phase of a fault. An added advantage of the yoke design is that no hoop stresses are developed in the windings because all modules are circumferentially independent. The analysis of the winding space for all but the torsionalloads is carried out in two steps. The first step is to analyze a single yoke and the superconductor inside of it. 142 K. A. Tepper, J. V. Miaerviai, 1111d J. L. Smith, Jr. Table U. Bending Stress in Yokes Locations* A 8 c D E F Bending stress, MPa 300 11 350 110 25 26 * See Fig. 2 for Iocations. The second is to model the winding space containing the superconductor filled yokes and the other rotor components. The analysis of the individual yokes will be given followed by the overall rotor analysis. One yoke and the superconductors inside the yoke are modeled as a frame filled with radial running strips of copper. The dimensions of the yokes are shown in Fig. 2. The copper strips, representing the superconductors, are the full height of the yoke by the width of one superconductor, approximately 0.15 cm. The rentrifugal and radial Lorentz forces push the superconductors against the yoke tops; however, since the yoke bottom and the bottoms of the superconductors are not bonded together the superconductors are in a state of compression. The maximum value of these stresses in the superconductor during the most severe operating conditions is u = -18 MPa. The azimuthat Lorentz forces are transmitted through adjacent layers of superconductors to the side walls of the yokes. Again, since the yoke side walls and the layers of superconductor are not bonded together there are no tensile stresses in the azimuthat direction. The compressive stress in the superconductor resulting from the accumulation of Ioad has a maximum value in the seventh module and is u = -20 MPa. The side walls of the yokes bendunder the azimuthat Lorentz force. The superconductors are bent into the shape taken by the side wall. The maximum values of the shear stress, r, and the normal stress, u, in the superconductors are u = 37 MPa and r = 3.9 MPa. In the second step of the stress analysis of the winding space, the yokes are modeled as frames in order to evaluate the stresses in the yokes resulting from the interaction of the winding space and the rest of the rotor. Since the superconductors are not potted into the yokes they do not contribute to the radial sti:ffness of the frames. This model has been coded into a multiple cylinder and truss program The peak stress Ievel is in the yokes closest to the junction of the two saddle coils and occurs during an armature short circuit when the radial Lorentz forces on the damper reach a peak. The results of this analysis are given in Table II. Two tests of the strength of the yokes have been made. The firstwas a test of the compressive strength of the yokes between the top and bottom bosses. The second testwas of the strength of the yoke 's side wall and top and bottom pin configurations. The results of the tests are in agreement with the predicted behavior. eJ. THERMAL ANALYSIS The yoke-module winding configuration is attractive for conductor cooling because it exposes a large fraction of conductor surface to the liquid helium while providing reserve helium capacity surrounding the modules. This gives good steady- Supereouductiog F1eld Winding for a 10-MVA Geuerator 143 state convective cooling and also allows for transient cooling by means of the large heat capacity of the liquid helium. The large power density of this generator requires a high-field winding-current density and this precludes the use of a completely cryostable conductor. The superconductor must always be maintained below the critical temperature. Any normal zone will quickly propagate throughout the winding in a rapid quench and thus local normal transitions must be avoided. A heat transferanalysiswas performed to determine the thermal response of the conductor modules to heat inputs into the winding space. The sources of these heat inputs are: conduction into the winding space from surrounding components (i.e., structural shells and electromagnetic shields), internal heat generation within the conductors, and frictional heating due to relative conductor motion. The internal heat generation is a result of eddy current and hysteresis Iosses in the composite superconductors when they are exposed to local alternating magnetic fields or during field current ramping. The values of the Iosses are calculated by a computerprogram that performs a dynamic simulation of fault transients on the generator including three-dimensional field effects [6 ]. The dual shielding arrangement of the generator provides excellent electromagnetic shielding of the field winding resulting in a peak ac Ioss rate of 1 mW/cm 3 • This Ioss rate was the peak average Ioss rate during a three-phase terminal fault from rated Ioad with the fault being cleared followed by reclosing to the bus. The total global winding loss rate for this condition was 8 W of which 6.5 W is from hysteresis and 1.5 W is from eddy currents. Although this is the peak Ioss rate, this value was used as a steady-state rate of heat generation to give a conservative estimate for the maximum conductor temperature. The frictional heating occurs at the interface of the slip planes as the azimuthal Lorentz forces increase during field current ramping. The mechanism is as follows. The increasing azimuthalload causes the side wall of the yokes to bend. The radial layers of superconductors bend to the same shape as the side wall. Since the interface shear is limited by friction, the layers deform as discrete beams, not as one thick beam. This results in relative motion across the slip planes. A worst-case analysis yields a peak heating rate of 0.6 mW/cm 3 if the field is charged in 30 sec. The heat transfer analysis consists of three thermal models: a two-dimensional steady-state model in a transverseplane (r-6}, a two-dimensional steady-state model in a radial plane (8--z ), and a transient adiabatic heating model. The two-dimensional model in the transverse plane assumes a uniform internal heat generation with no axial conduction. The two-dimensional model in the 8--z plane assumes uniform internal heat generation with no radial conduction. Both models utilize the effective conductivities of the conductor-insulation composite cross section with surface heat flux through convective heat transfer coefficients. The maximum conductor temperature rise for an internal energy generation of 1 mW/cm3 is 0.01 K. The transient adiabatic temperature rise resulting from a uniform heat input to the winding space was estimated by considering the heat capacity of the helium. This temperature rise during a fault is 0.3 K. The winding will remain superconducting under all operating conditions because the maximum temperature will not exceed the critical temperature (approximately 5 K) foranominal operating temperature of 4.2K. CONCLUSIONS A prototype electrical generator of 10-MVA rating with a superconducting field winding is being constructed. A unique yoke-module type of support system is being 144 K. A. Tepper, J. V. Milleniai, .... J. L s.ltll, Jr. employed in the field winding. Structural and thermal analyses of this configuration indicate that this system will provide stable mechanical, thermal, and electrical operation under both steady-state and transient conditions. REFERENCES 1. J. L. Smith, Jr., G. L. Wilson, and J. L. Kirtley, Jr., IEEE Trans. Magn. Mac-15(1):727 (1979). 2. J. L. Smith, Jr., G. L. Wilson, J. L. Kirtley, Jr., and T. A. Keim, IEEE Trans. Magn. Mq-13(1):751 (1979). 3. "Demonstration of an Advanced Superoonducting Generator," Interim Rept. 7, MIT Cryogenic Engineering Labaratory and Electric Power Systems Engineering Labaratory, E(49-18)-2295-7 .Oll, IR 7, pp. 3-5. 4. J. L. Smith, Jr ., J. L. Kirtley, Jr ., P. Thullen, and H. H. Woodson, Proc. 1972 Applied Superconducßvity Conference, IEEE Pub. No. 72, CH0652-5-TABS (1972) p. 145. 5. "Demonstration of an Advanced Superoonducting Generator," Interim Rept. 7, MIT Cryogenic Engineering Labaratory and Electric Power Systems Engineering Labaratory, E(49-18)-2295-7 .Oll, IR 7, pp. 232-238. 6. S. D. Umans, P. B. Roemer, J. A. Mallick, and J. L. Wilson, "Three Dimensional Transient Analysis of Superconducting Generators," presented at IEEE Power Engineering Society, 1979 Winter Meeting. C-3 CONDUCTIVE ARMATURE SIDELDING DESIGN CONCEPTS FOR SLOW-SPEED SUPERCONDUCTING GENERATORS IN THE 40- TO 400-MVA RANGE S. Kuznetsov Imperial College of Science and Technology London, England INTRODUCfiON This study is concerned with the electrical dimensioning of passive, electromagnetic shields surrounding air gap armature windings for slow-speed multipolar superconducting-field generators which make exclusive use of conductive rather than integrated conductive and ferromagnetic shields. Special reference is made to the problern of screening the composite stator winding and superconducting field for environmental and improved short-circuit protection rather than being applied to damp transient rotor oscillations. The conceptual machines used to illustrate the limitations on exclusive conductive shielding are 6-48 pole units with a corresponding stator rating rauging from 40 to 400 MVA at 60-Hz output. The specific cryogenic rotor circuit incorporates a Nb-Ti superconductor cable with design details included to calculate winding inductances and intermagnet forces. A high-voltage multilayer stator winding is dimensioned for numerically determining the Maxwell stress vector in the vicinity of the stator shield and reducing the load-dependent, radial magnetic forces on the superconductor. The motivation for proposing the construction of !arge slow-speed superconducting generators as peaking units for electric utility applications and as central drives for ac ship propulsion is based on the following. 1. The basic geometry of the slow-speed generator coupled with the use of conductive shielding indicates that the machine is fundamentally much stifler than an equivalently sized generator with ferromagnetic stator shielding. The increased no-load eddy current Iosses in the conductive shield are partially oflset by the substantial reduction in reactive kV A attributable to stator leakage reactance, while the rotor-to-stator mutual coupling remains virtually unchanged. 2. The conductive shielding concept reduces the armature reaction eflect of the outermost winding layers compared with alternative screening methods allowing better voltage regulation and armature conductor mechanical protection. 3. The conductor stator shield is clearly superior over a ferromagnetic shield in limiting transient forces on the superconducting rotor during stator short circuit for slow-speed machines where the synchronous and transient reactance are typically under 10%. 145 146 S. Kuznetsov Apart from the greatly increased liquid helium refrigeration requirements encountered with multipolar machines over high-speed machines with an equivalent field MMF rating, the major constraint on the use of high-field type II superconductors on the rotor is the steady-state azimuthalload torques and radial magnetic push which exceed those experienced in superconducting turbogenerator designs. A practical upper Iimit on the realization of slow-speed superconducting generators (SSSG) using Nb-Ti field coils was depicted by the design of a 400-MV A-size, 150 rpm hydrogenerator based on an equivalent rotor peripheral speed, active height, and pole pitch as a conventional615-MVA hydrogenerator [ 1]. The range of design considered in this analysis all retain the same rotor peripheral speed of 70 m/s ) l!llildjll ,11111111 P @) Qll!llulll!jdj! dll ffil!llll ill!illll' b ii!HHHHHiffllllll ill!ll!!lll'llllllllll lllllllilil!ll lllilb cllllillllllllllllllll IIIIIIIIIIIIIIIIIIIIP illllilll!,jjjii!IIIP cllllililillllllllllll llliilllllilllllllllb qll!lllilillllllllill jll!!lillll!l!lll!tll llll!lllilll!Jl!!I!!D llilllllll!iilll'lllb lillllilll!l!ijj,II!P al!rtll!!llllllll!lll llllilljljliljllilllb ~llllillllllllllll lllllllllllllllll~ qllllll!llllllllll lllllllillllll l!l~ Airgap 0 Fig. 1. One-half-pole segment radial-azimuthat view of conceptual SSSG machine. Legend: (0) aluminum armature shield; (1) EHV stator Roebel conductors; (2) stainless steel rotor damper shields; (3) copper ambient temperature damper; (4) compression spacers; (5) vacuum; (6) 80-K heat radiation shield; (7) 4.2-K liquid helium COntainment vessel; (8) rotor torque transfer banding; (9) Nb-Ti superconductor field winding; (10) liquid helium cooling ducts; (11) field winding COntainment block; and (12) rotor torque axial transfer coupling. Conductive Armature Shielding for Slow-Speed Supereonducting Generators 147 Table I. Comparative Dimensions and Performance of Representative Slow-Speed Superconducting Generators 6-Pole unit 48-Pole unit Rotor Diameter, m Active height, m Weight w/o shaft, kg Peripheral speed, m/s WR2 , kg-m 2 Inertia constant (w /o shaft) Mass of superconducting cable, kg Cost of field winding conductor, 1979 $ Nominal ac loss in superconductor, kW Field array Iead refrigeration, kW Torque transfer refrigeration, kW 1.11 3.45 7.7 X 103 70 2.12 X 10 3 0.419 1.54 X 103 3.3 X 105 2 3 190 8.92 3.45 5.4 X 104 70 9.6 X 105 0.30 1.08 X 104 2.3 X 106 16 3 1700 Stator Diameter, m Mass, kg* Rating, 60°C rise, MVA Terminal voltage, kV Air-gap, mechanical, mm Acti_ve height, m Current loading, A/m periphery Short circuit ratio Load 1 2 R loss, kW No-load eddy current loss, kW Efficiency at 0.90 PF Specific power density, kVA/kg 1.47 4.1 X 10 3 40 20 19 3.88 34,000 >11 188 235 96.6 2.55 9.3 3.56 X 104 400 230 19 4.02 36,000 >15 1140 1520 98.8 3.12 * Winding and shield only. and rated rotor Ioad torque of 480 kN-m/pole acting on a pole area of 1.83 m2 • A segment of the 400-MVA machine taken in the radial-azimuthat plane is shown in Fig. 1; note that the radial curvature is very small. The dimensions and performance specifications are included in Table I. The major areas of concern are: (1) the critical thickness of the conductive shield in a solid-shell form [2 ] that depicts the transition from inductance-limited screening to resistance-limited screening; (2) matehing the reftected surface impedance of the conductive screen to the combined reaction wave of the field and stator MMF's and establishing the optimum degree of inductance-limited screening as a function of rotor Ioad angle; (3) maintaining the external magnetic field to less than 0.5 mT at a distance of 2m from the stator shell so that these machines are adequately screened for ship propulsion applications; (4) examining the confticting electromagnetic requirements demanded by the dual use of a single conductive shield for both steady-state and transient operation; and (5) the relative performance of the stationary conductor stator shield against experimental evidence on rotating warm damper shields (excluding inertial free shields) in limiting the specific transient radial magnetic stress applied to the superconducting cables under terminal short circuit. Although the rotor heat radiation and warm shields are shown in Fig. 1, their effect is not included in the foregoing magnetic field plots so as to isolate the singular effect of the stationary-shield concept. 148 S. Kuznetsov ROTOR FJELD STRUCTURE In contrast to conventional superconducting turbogenerator rotor designs whereby all of the rotor structure from the cold radiation shield inwards is maintained at a cryogenic temperature eJ, the cryogenic portion of the described slow-speed superconducting rotor is constrained to a relatively narrow band of about 0.14 m in radial thickness and mounted on an ambient temperature rotor spider to connect with the central shaft. Most important, the high-torque SSSG is only practical when the field coils are abnormally sieoder in comparison with turbogenerator field coils. For the reference design, all of the superconducting coils have a width to pole pitch aspect ratio of 6.46 : 1. This shape is avoided in single isolated high-field, air-core coils as the Lorentz forces tend to ovalize any quasireetangular shapes. However, when a series of long "racetrack"-shaped superconducting coils are situated in a tightly packed band along the rotor spider periphery, a uniformly stressed array can be obtained for the SSSG at a rotor radius of curvature ranging from 0.55 to 4.46 m. Fundamentally, high-aspect magnets are only permissible in rotating machines when their shortest transverse dimension is aligned in the same direction as the rotating magnetic field in order to avoid any large differences in tangential torque applied to the leading and trailing conductor portions of the coils. The advantage gained by going to a large multipolar SSSG design is brought out by the apparent enhancement of the superconductor's criucal current capability simply by the reduction in internal magnetic field for a given air-gap induction density. For example6 at rated Ioad, the field MMF of the 400-MVA machine amounts to 0.54 x 10 A-turns/pole. If this value of excitation is applied to a single isolated magnet, the Nb-Ti critical current density would be based on a peak llcm __I____ _ '' ' \ \ \ I \ I I I I I I I I I I I I I / ... - ...... \ \ \ I I I I I \ I I I I / / I I I to rotor shar' I I Fig. 2. Contour lines of constant radial magnetic flux density (in mT) of an axial-radial segment, 16.7° from the direct axis. Conductive Armsture Shielding for Slow-Speed Superconducting Generators 149 magnetic field of 3.3 T. With similar magnets arranged immediately adjacent (and oppositely polarized) in an array, the peak internal field is 5.3 T. Specifically, at a pole pitch of 0.57 m, the use of a !arge radius, multiple magnet system results in an 18% reduction in critical internal field intensity per ampere of excitation. ROTOR FIELD REPRESENTATION The magnetic field and Maxwell stress vectors (Fig. 2) are calculated by performing a discrete conductor, three-dimensional integration over all of the conductors contained in both the rotor and stator assernblies using a variation of the Biot-Savart law in a closed-form solution. The crucial numerical technique isthat all coil shapes must be discretized into exclusively orthogonal components since there is no mutual coupling between components in different planes. The orthogonal components of the model are made sufficiently smaller than the cable cross section; both the axial and radial curvature of the superconducting coils including the end regions are represented in the discretization with an error not exceeding 1 part in 103 for the total field to stator mutual coupling. A flux plot of the axial component of magnetic field density for the 400-MVA machine at a radial distance of 0.13 m from the field pole planar center line is shown in Fig. 3. In this one specific case, the armature conductor active length is only marginally Ionger than the superconducting field coil. This gives instant visual representation of the undesirable radial magnetic forces acting on the first layer of armature winding. Fig. 3. Contour lines of constant axial magnetic ftux density (in mT) in an axialazimuthat segment, 13-cm radial from the field pole face at the first stator winding layer. (Shortened axial active length of 0.80 m, pole pitch of 0.57 m, and field excitation = 540 kA-turns/pole). 150 S. Kuznetsov STATOR ELECTRICAL DESIGN To appreciate the most general effect of stator conductive shielding, a nonhelical, nonspiral concentric wave winding with straight conductors has been used in the conductor discretization. The stator pole pitch progressively increases as a function of radial distance and an azimuthat offset in individual conductor spacing amounting to a 4° shift/layer has been employed for the primary purpose of reducing peak electric field potentials and for valtage harmonic control. In this most basic stator structure, the end region connections are nearly free of axial magnetic stresses by having the active length of the stator exceed the active rotor field length by 0.55 m per side (or 16% oversizing). This increase in axiallength yields a 2.7% increase in fundamental rotor-stator coupling and reduces the second harmonic component of axial force on the stator layer nearest the air gap from 7.9 to 0.6% using the rated tangential force on the SSSG as a base. CONDUCTIVE SHIELD SELECTION Ideally an electromagnetic shield surrounding the stator winding should be a composite, Iaminated ferromagnetic and conductive structure to fully satisfy the differing steady-state and transient requirements [4 ]. However, a single solid aluminum shell clearly identifies its potential inftuence on limiting field and stator coil mechanical stresses. If stray eddy current Iosses were of no concern, then it would be desirable to individually shield each concentric layer of stator winding. As a compromise between minimizing capital investment and short circuit rotor forces vs. minimizing stray eddy lasses, an upper bound on the maximum shield dissipated power at no Ioad has been established at 100% of the stator rated Ioad 1 2 R lass (based on a stator current density of 3 A/mm 2 ). The inductive coupling between each armature layer and the shield is shown in Fig. 4. To ensure that the inclusion of the stator shield can be accurately represented in the magnetic field analysis of the entire machine, it has been necessary to discretize the solid conductive shield into finite conductor sections with judicious care taken to define the shape of the eddy current paths as would occur on the real machine. The ratio of the axial current density to the azimuthat component of current density in the shield is shown in Fig. 5. The bulk of the experimental work carried out on air-gap winding machines of this type has been directed towards plotting the exact eddy current distribution in continuous shells as a function of transient time constants, reftected impedances, and axial overhang lengths. These tests have concluded that in the recommended inductance-limited shells for the range of 40- to 400-MVA machines, all shields have nearly perfect one-dimensional axial current ftow in the portion covering the rotor pole face. This experimental evidence has obviated the need to slit the conductive shield axially at regular intervals to ensure unidirectional ftow under the transients originally envisioned. All of the computational results have assumed that the conductive shell is composed of standard 2219-T81 aluminum ofresistivity 3.7 x 10-8 0-m. In general, the shell thickness of 0.014 m is near optimum based on the universal, 60-Hz shield power transfer curves of Fig. 6. The radial attenuation of the shield reaction field in Fig. 3 is approximately 160 mT /m; at the first or bare winding layer, the leakage reactance is only changed from 2.55 H/m(axial)/m (periphery) to 2.23 H/m/m. The effect of the single stator shield has changed the machine synchronaus reactance Conductive Armature Shielding for Slow-Speed Superconducting Generators 151 40~~--~----~---,----,---~----~--~----,-~ 6 _; 0 ~ E i 30 "'- ., 6 400 MVA Design 0 10 0 MVA 6 0 6 V c: 0 :! V :> 6 0 ." -" ." c: ~20 0 u 6 0 6 0 "'c: ·~ c: ~ ~ u;"' 0 ;; 10 -.; ·~ "' 0 ~~12~--~--~16----~--~20~--~--~ 24~--~--~2~8~ Rodiol Seporat ion belween Shoeld and Stator Wond•ng Layers,cm Fig. 4. Shield to stator winding coupling inductance for 0.014-m-thick shield and two-conductors/pole/phase stator with end region conductors included. from 0.078 to 0.066 pu. For these machines, the total stator current loading is 36,000 A rms/m periphery while the equivalent shield current loading is 7000 A/m for a magnetic density at the shield inner surface of 870 mT. In terms of the shield's direct effect on the rotor torque, an additionalloading of 1.51 kN-m perpole is incurred at rated Ioad in the 400-MVA design and approaches 2.0 kN-m/pole in the 40-MV A type. The balanced-phase, centered-rotor radial magnetic pull on the superconductor integrates to 4.8 x 104 N/pole at rated Ioad and 0.90 power factor leading owing to the stator winding alone; this increases to 5.7 x 104 N/pole with the addition of the shield. These two criteria of rotor torque and radial magnetic pull are used exclusively to evaluate the performance of the shield because they independently express the energy stored in the direct and quadrature axes of the generator. For example, if the SSSG is Operated in the lagging power factor mode (absorbing reactive power) at about 20% of rated stator loading, then it is possible to effectively neutralize the radial magnetic force on the superconducting field magnets for the particular inductance-limited shell recommended; this effect has been repeatedly demonstrated in laboratory machines. 152 S. Kumetsov I0 2r------,-------r------,-------,-------1, -- -- - - , ... .t: :; .....E .., 10° .. .. .~ M .., .':! 0:: 1Ö1 d=0.01' m 60 Hz 1>= 3.7x 10·8 o.s 1.0 Xar ia l 1.5 per uni t Fig. 5. Mean ratio of axial current density to azimuthat current density in stator shield as a function of axial distance from the machine centerline per unit of stator conductor active length per side. LABORATORY INDUCTION SIMULATOR A full-scale conductive shield of the type described and applicable for an 80-MVA SSSG is currently undergoing laboratory tests in the form of a 2.2-mdiameter rotating disk annulus with a high-field, 80,000-A/m air gap, axial ftux stationary excitation winding to simulate the predicted magnetic field density at the shielding interface experienced in the superconducting machines. The experimental facility is shown in Fig. 7 with the Maxwell stress instrumentation in the air gap and telemetry of the rotor disk current ftow distribution. Aluminum, Dural, and solid steel disks of up to 30-mm thickness have been evaluated up to tip speeds of 140 m/s. The time-transient diffusion effects in shields as !arge as those proposed for the 400-MVA unit have been simulated by the use of an azimuthal-propagating space Conductive Armature Shielding for Slow-Speed Superconducting Generaton 153 2 ' 1 C 'r+-~-------.-------.--------.-----, N E ~ ...c,.. .. .. 104 0 .3 ~ ~ 0 0.. .., -.; &. V'l V .. V a. "' 5 110 3~......__ _ _ _ __.__ __ _ ____._ _ __ 5 10 Shield Thickness, m m _ _.___....J 20 15 Fig. 6. Specific shield power loss density for positive sequence fields at 60Hz and constant field current of 500 kA-turns. edge effect produced by a discontinuous, step-function-excitation winding arc of 180° acting on a continuous rotor. Under the condition that the shield is to be operated in the inductance-limited mode, which is valid for general line-induced transients in the generator, the eddy current loss in the full-scale shield may be directly calculated from measurements made from only the azimuthat component of the magnetic field intensity at the model shield inner periphery. Furthermore, when the induction simulator is operated at synchronous speed, the azimuthat buildup of the air-gap magnetic ftux density traveling wave actingintime quadrature with the excitation current may be expressed [5 ], without recourse to numerical solution as _ #Lo ~ T;lsi(l bq (t) - L.. 1T i~l g; - e -X?r/T·wT.) · · cos ( wt - X ) -7r T; (1) where 2 T; = T~#J-0 1T pg; ( 2) and x is the azimuthat distance under the constant 1. excitation winding that an eiemental rotor segment has moved since entry at hq = 0. The magnitude of bq(x) at a 154 S. Kuznetsov Fig. 7. High-speed Iaboratory induction simulator. Axialftux, air ~tan winding of 180° arc excitation to simulate transient shield operation by use of azimuthal magnetic diffusion edge effect. point x obtained from (1) is related to the instantaneous value of bd(t') in the full-scale salient pole generator at a time t' from the initiation of the actual time transient at t0 as TW x = - ( t ' - t0 ) (3) 1T In the laboratory induction simulator, with a 2.2-m aluminum annulus, 9.5 mm thick, and resistivity of 3.25 x 10- 8 0-m, a shield magnetization time constant, T, of 89 ms has been attained for a pole pitch of 0.253 m. The advantage of this experimental apparatus is that with this particular disk, shields for a 60-Hz superconducting generator up to 19.4 mm thick may be evaluated by frequency scaling to 250 Hz in tests, in addition to using a continuously occurring space transient to simulate rapid, time-rlependent eddy current diffusion effects. CONCLUSIONS The advantage of the passive, conductive armature shield is its capability of reducing steady-state magnetic forces on air gap armature windings, and in contrast Conductive Armsture Sbielding for Slow-Speed Superconducting Generators ISS to saturable ferromagnetic shields, the conductive shield provides better protection of transient forces on the field superconductor and rotor damper shield during symmetrical and negative-sequence faults. lts disadvantages are increased steadystate power dissipation, larger balanced rotor torques, and lower values of synchronous and subtransient generator reactances. Additionally, the stator conductive shield reduces the radial magnetic forces on the field superconductors in steady-state operation when the generator absorbs reactive power. The multipolar superconducting machine is thus practical either as a utility peaking unit or as a ship propulsion motor where the power-to-weight criterion takes precedence over absolute efficiency. ACKNOWLEDGMENTS The author would like to express bis gratitude to E. R. Laithwaite of the Heavy Electrical Engineering Department at Imperial College for his technical advice, and to the National Research Development Corporation of England for the use of their laboratory machinery. NOTATION b = instantaneous magnetic flux density g, = airgap between ith stator winding and shield inner periphery J., = stator winding current loading of ith layer t =time T = stator shield magnetization time constant x = azimuthal distance from entry edge of excitation winding X = reactance at 60Hz z = total nurober of stator winding layers Greek Symbols p. 0 = free-space permeability w = electrical radian frequency p = resistivity T = pole pitch Subscripts d = direct axis i = stator winding ith layer q = time quadrature component s = stator REFERENCES 1. S. Kuznetsov, IEEE Trans. Magn. Mag-1S:719 (1979). 2. G. G. Lessmann, W. A. Logsdon, R. Kossowsky, M. P. Mathur, and J. M. Wells, "Structural Materials for Cryogenic Applications," Rept. 74-904-CRYMT-2, National Bureau of Standards Contract No. CST-9304, Westinghouse Electrical Corporation, Pittsburgh, Pennsylvania (1974). 3. P. Thullen, T. A. Keim, and J. V. Minervini, IEEE Trans. Magn. Mag-11:653 (1975). 4. C. J. Carpenter and M. Djurovic, Proc. IEEE 122:681 (1975). 5. E. R. Laithwaite, Induction Machines for Special Purposes, Chemical Publishing Company, New York (1966), p. 76. C-4 OPTIMIZATION OF SUPERCONDUCTI NG CRYOTURBOGEN ERATOR FJELD-WINDING PARAMETERS B. I. Verkin, I. S. Zhitomirskü, and R. V. Gavrilov Physico- Technical Institute of Low Temperatures Academy of Seiences of the Ukrainian SSR Kharkov, USSR INTRODUCTION A mathematical formulation and algorithm is proposed for the design of superconducting cryoturbogenerator field windings, e.g., saddle-shaped windings. The choice of the particular set of component dimensions in a cross-sectional plane of a rotor can be optimized in combination with the dimensions of the supporting elements for the winding and cooling channels as weil as the percentage of superconducting material in a cable cross section. This set of variable parameters is considered to be optimum for a given external diameter of a winding if it results in a maximum ftux in the armature-winding region (corresponding to the unit length of an active part of the winding) and it meets the constraints of cryostatic stability and strength. Examples of computer utilization of the proposed algorithm are given. Results show that the algorithm is very effective. PROBLEM AND METROD OF SOLUI10N The objective function that is normally chosen in the optimal design of superconducting cryoturbogenerator field windings is the magnetic ftux <I> averaged over the radius of an armature winding for unit length of active winding. Providing the maximum <I> makes it possible to attain the maximum specific power by means of only n parameters which define the superconducting field-winding configuration in a rotor cross section. Substitution of Z for this n-dimensional vector and the superconducting current density by j gives the relation <I> = <I>(Z, j) = <I>(Z, 1}j (1) In order to obtain the optimum configuration Z and current density j we must consider the following basic requirements: 1. 2. j ::5 ß min ico with the minimum being taken over the values n at different points in the winding. The mechanical stresses resulting from the centrifugal and electrodynamical forces in the superconducting field winding and its supporting elements must not 156 Optimization of Superconducting Cryoturbogenerator Field-Winding Parameters 3. 157 exceed the values permissible for a given material at a specified loading and temperature range. The requirements as to shape of the magnetic field generated by the superconducting field winding in the armature-winding region must be satisfied. Using the constraints of 1 and 2 above tagether with the inequality (1) makes it possible to eliminate the variable j from the adjustable parameters. Restrietion 3 is best considered by means of a penalty function Cl The problern then reduces to a search for an absolute extremum. To solve the problern posed above the method of configurations has been used [ 2 ]. The calculation of the magnetic field and electrodynamic forces was made with a method which takes account of the discrete winding distribution eJ. EXAMPLE: WINDING IN PARALLEL SLOTS An application of this method is the optimal design of the winding for the cross-sectional scheme presented in Figs. 1 and 2. Here the winding is embedded in parallel slots of the monolithic rotor. The centrifugal and electrodynamic forces affecting the superconducting field winding are transferred to the teeth (1) and wedges (2); these must satisfy the strength requirements described previously. In addition, the strength requirements of the compounded winding must also be satisfied. At a given external rotor radius r, the geometry of the winding cross section is uniquely defined by the parameters y;, a;, b;, (i = 1, 2, ... , v). These parameters form the vector Z and are also the unknowns in the design. 1 3 X Fig. 1. Schematic view of cryoturbogenerator cross section. Legend: 1, superconducting field winding; 2, armature winding; and 3, ferromagnetic shield. 158 B.I. Vedda, I. S. Zldtominldi, md R. V. Gavrilov y 4 X Fig. 2. Schematic of rotor cross section. Legend: 1, framework; 2, superconducting field winding; 3, framework teeth; and 4, wedges. I. +--- Fig. 3. Optimized geometry of rotor cross section (dashed line) compared with geometry estimated before optimization (solid line). Optimization of Snpercondneting Cryotarbogenerator Field·W"mding Pu81Deters 159 Table I. Winding Parameters Magnetic llux, Wb/m S/Cu fieldwinding area, m 2 Numberof turns perpole Permissihle current, A 1.07 1.18 0.0294 0.0275 1820 1700 1000 980 Zo Z* Figure 3 shows a comparison of the calculated optimum vector Z* obtained from the above-mentioned method (dashed line) and the geometry corresponding to the initial vector Z 0 derived from preliminary considerations (solid line). lt follows from Table I (where some parameters of the above windings are given) that witp this optimization scheme the magnetic fl.ux was increased by 10% and the superconductor requirement was reduced by 7%. Note that among the constraints given, only the first and the second turned out to be operational ones since for these equality is attained. With respect to constraint (3), it was satisfied over the entire range investigated. OPTIMIZATION OF TUE CRYOSTABILIZATION PARAMETERS In the large cryoturbogenerators currently projected for utility networks, the stabilization problern of superconducting field windings assumes ever greater importance particularly since reliability is so critical. Load asymmetry and its time variations, vibrations and hunting oscillations of the rotor, and helium fl.ow rate variations Iead to disturbances of the magnetic field, exciting current, and temperature. These disturbances can be sufliciently large to result in quenching. Shutting off a generator in an emergency when the quenching takes place is unacceptable because it causes surges in a utility network. The only acceptable solution to the problern is the attainment of cryostatic stabilization which permits the winding to return to the superconducting state from a full or partial quenching. In view of these design considerations, a scheme which uses transverse cooling (i.e., one with the helium channels normal to the winding conductors) and where each conductor appears tobe subjected to alternating cooled and noncooled regions is more useful for the large cryoturbogenerators (Fig. 4). Let us first consider that part of the conductor between the cooled and noncooled regions. lf the lengths of the cooled and noncooled regions, 211 and 21, 1 I 2 ~~ " 1 \ ) II I I I ' I X 0 Fig. 4. Scheme for cooling of winding. Legend: 1, monolithic layer of superconducting field . winding; and 2, helium channels. 160 8. L Veftm, L S. ZW...UIIIdi, IUid R. V. Gamlov respectively, are much greater than the characteristic dimensions of the conductor cross section, then the conductor temperature field may be considered as one dimensional; i.e., T = T(x). This temperature field can be described by a nonlinear heat-conduction equation T; the nonlinearity is a result of the temperature dependence of the heat capacity C.., heat conduction A.., and electrical resistivity Pe· To simplify this procedure the equation may be linearized so that the solution f to the linell! pro_blem would be a "worst-case" solution of the original nonlinear problem, i.e., T s T throughout the range of interest. The worst-case conditions are obtained, for example, by substituting the maximum value PM for Pe in the temperature range considered and by substituting the minimum value AM for Ae. For a stationary normal zone, assumed to be in a conductor and occupying the region 0 s x s x 0 , there is some value for the helium heat-transfer coeffi.cient a 0 where a > a 0 that permits the problern with the above boundary conditions to reduce to 0 <X< I+ l1 dT =O dx ' X= 0, X = I+ l1 (2) where [ 2 QM= { PMA2' 0:5 X :5 Xo 0, 0, { OHe = a(To- T)P A ' and has a solution I$ X$ I+ 11 f satisfying the condition T(x 0 ) < Tc(I/ {A, B) (3) This means that the critical nonlinear problern has a stationary solution in the normal zone which is less than that expected from (3). Now Iet the values of a 0 , corresponding to different x 0 , be limited and a * = maxosxosl+h a 0 • Then if a > a*, a normal zone of any size will return to the superconducting state and the coil will be completely stable. As an analysis of (2) shows, the condition of complete stabilization may be represented as follows: (4) where s must be less than unity. lt may be shown for given values of B and T0 which satisfy (4) that the maximum value of the average current density I/ A may be obtained when the fraction of superconductor in a cable cross section is given byt = [o.5 + (0.25 + K-1/2>1/2r1 (5) r t Results similar to these have been obtained by Wipf [4 ). Opdmization of Sapercondnc:ting Cryotarbogenentor Field-Winding Pllftllleters 161 where K = aPI1 ic(To, B)(iJic/iJT)PcuA(l + lt) The maximum avaiJable current density which meets the stabilization requirements is obtainable from the relations isc =(*je (6a) I/A = (* 2jc (6b) Considering that, as a rule, K « 1 we may write that r* = K112, (I/A) !> =K112. = [ max ]c aPltic(To, B) ] 112 A(l + lt)PculiJic/ iJTI Satisfying the requirements for cryostabilization at a current density which is not too small appears to be possible if one takes care of the following: 1. The field winding should be fabricated of a material having a ratio of ic/liJic/iJTI as large as possible (for example, Nb 3 Sn). 2. The structural design should be made to provide suffi.ciently effective heat exchange between the helium and the winding conductors. As evaluations have shown, there exists a real possibility of obtaining a value of more than 106 W /m 3 K for the complex aPld A(l + 11 ). To account for the effect of induction and helium temperature changes, B and T0 , respectively, over the winding thickness it is necessary to solve a much more complex problem. For such a design scheme of a superconducting field winding and outer rotor radius we must determine the superconducting field winding current density, the distribution of the winding throughout the rotor cross section, the dimensions of the channels through which the helium is circulated, the fraction of superconductor in the cross section of the superconducting cable, and the dimensions of the supporting elements at which the magnetic flux of an exciting field is maximum with respect to the active portion of the cryoturbogenerator. In addition to requirements (1), (2), and (3), noted earlier, we must take into account the requirement of cryostatic stabilization for the superconducting field winding. NOTATION A = cross-sectional area of conductor a; = thickness of ith tooth b; = width of ith modulus B = magnetic induction I = transport current j = current density critical current density current density in superconductor P = perimeter of cooled surface T0 = helium temperature Tc(j, B) = critical temperature of superconductor carrying current of density j at a ftux B Y; = height of ith modulus ß = safety factor allowing for the possible degradation, ß < 1 ( = superconductor fraction of a cable cross section 211 = number of moduli forapole Pcu = resistivity of copper matrix at 20 K cl» = magnetic ftux ic = isc = 162 B. I. Verkin, I. S. Zhitominkii, and R. V. Gavrilov REFERENCES 1. E. Polak, Computational Metlwds in Optimization, Academic Press, New York (1971). 2. D. J. Wilde, Optimum Seeking Metlwds, Prentice-Hall, Inc., Englewood Qiffs, New Jersey (1964). 3. I. S. Zhitomirskii and R. V. Gavrilov, Magnitnoje pole i elektrodinamicheskije usiliya v ekranirovannoy elektricheskoy machine s nemagnitnym serdechnikom, Izvestiya AN USSR, Energetika i transport (5):122 (1977). 4. S. L. Wipf, "Stability and Degradation of Superconducting Current-Carrying Devices," Los Alamos Scientific Laboratory Rept. LA-7275 (December 1978). D-1 SUPERCONDUCTING COMPENSATING SOLENOIDS FOR THE CELLO DETECTOR EXPERIMENT AT PETRA W. Barth, N. Fessler-Wilhelmi, W. Lehmann, and P. Turowski Kernforschungszentrum Karlsruhe Institut für Technische Physik Karlsruhe, Federal Republic of Germany INTRODUCTION One of the main components of the detector experiment CELLO in the e + estorage ring PETRA at DESY /Hamburg is the superconducting magnet system. The magnets have been developed as a joint project by CEN/Saclay and Kernforschungszentrum Karlsruhe. CEN has built the thin-walled main solenoid with a length of 3.5 m and a winding diameter of 1.7 m creating a magnetic field of 1.5 Tin the center CJ. Two magnetically identical solenoids, located symmetrical to the main coil, must be positioned with a magnetic field opposite to the central field to cancel the integral field effect on the orbiting particle beams. This compensating solenoid system, including its supporting cryogenic system, has been designed and constructed by the Kernforschungszentrum Karlsruhe. The main coil was constructed with a wall thickness of :s.A/2 (equivalent to 45 mm of aluminum); this requirement was imposed by the radiation length A of the generated particles. The compensating coils, on the other band, were designed and built by conventional techniques for superconducting magnets. The solenoids within their cryostats are to be mounted in movable iron yoke doors surrounded by muon chambers and a liquid argon end cap calorimeter as shown in Fig. 1. The outer geometry and the required warm bare parameters determine the dimensions of the cryostats, i.e., a length of 1.10 m and an outer and inner diameter of the horizontal part of 0.8 and 0.4 m, respectively. The compensating coils are parts of a 4.2-K helium circuit which includes the main solenoid (cooled by forced-flow helium), a complex helium transfer line system and a refrigerator, and have been built by CEN/Saclay. The compensating coils are cooled by a liquid helium bath and the cryostats have shields cooled by the return gas. Each cryostat is connected in parallel to the refrigerator over a length of 13 and 18 m, respectively. The design of the magnets and cryostats along with their peripheral equipment and the cooling flow scheme is reported here. Results of the tests on the completed equipment including magnet performance, thermal insulation, safety aspects, and the status of the final installation at DESYare also presented. 163 1 2 J 4 5 6 7 8 9 10 11 12 13 Vacuum Beam Pipe Superconduding Coil of Oetector fron Yoke Compensation Coils Movmg Devices Feed Llnes lor liquid Helium MPI,Nünchtn 101. brlwht OESY. Hllllburg Unin nity IiD ol Pari• S.ctoy Orur ParticiP.anls : Fig. 1. Schematic view of the detector experiment CELLO (DESY). Centrat Orilt ·and Proportional Chambers Endcap Proportional thambers Endcap Shower Counters Iliquid Argon 1 Cylinlfric Shower CountersUiquid Argon] Proportional Chambers for MuOn Detecfion Drift Chambers tur Forward Detedor Shower Counler for Forward Detector <~1:1-1-<» Total Weight : ~ 1400 I Magnet Aeld ' 15 kr DETECTOR I :-= ! ~ i ~ l ~ ~ ?'! j ~ :'"" Superconducting Compensating Solenoids for the CELLO Detector Experiment 165 Table I. Solenoid and Superconductor Parameters Solenoid Length of coil, m Inner winding diameter, m Outer winding diameter, m Centrat field, T Maximum field at conductor, T Nominal current, A Overall current density, kA/cm 2 Max. stored energy with iron yoke, MJ Max. stored energy as air coil, MJ Superconductor Cross section without insulation, mm Filaments, ,_.m <P Twist, mm Cu/SC ratio Critical current at Bmax of 3.8 T, A Residual resistivity ratio of Cu Insulation 0.86 0.5 0.54 3.5 3.8 1050 15 -1.4 1 3.1 X 1.9 460 X 50 50 5: 1 2380 -280 Wrapped Capton foil with a thickness of 50 ,_.m COIL DESIGN AND CONSTRUCTION The important parameters of the two identical solenoids mounted within the iron yoke are given in Table I. The superconductor consists of a multifilamentary NbTi wire of reetangular cross section manufactured by Vacuumschmelze Hanau. The parameters of this conductor are also listed in Table I. Short sample measurements with this superconductor were performed by the transformer method [2 ]. The results are shown in Fig. 2 together with the Ioad line of the magnet. A stainless steel tube forms the inner part of the liquid helium cryostat and also acts as the coil mandrel. The mandrei was supported within the bore during winding by a holding fixture to take the reaction forces from winding the conductor estimated at 180 N. The conductor has been wound in ten layers with interlayers of glass rows. Each layer was impregnated during winding with epoxy resin Araldit CY 221 and hardener HY 979 supplied by CIBA. The flanges of the coil former were prepared with a nonadhesive Teflon foil to prevent the release of crack energy between the winding and the flanges during the magnetic field rise. The superconductor was fabricated in two lengths for each coil. The ends of the conductors were connected by an uHrasonie welding technique where the contact resistance is lower than 10- 8 n. eJ CURRENT LEADSFORmE SOLENOIDS Current to the compensating solenoids is supplied by gas-cooled Ieads. The CELLO experiment requires that there should not be any atmospheric condensation on the helium return gas lines and no mist formation on the head of the Ieads. Accordingly, the return gas lines are guided within the vacuum jacket of the liquid helium transfer lines while the room temperature ends of the Ieads are warmed by heater elements. The inner part of the current Ieads consists of an OFHC-copper braid whose Jength and cross section were optimized by using the calculations of Lock [4 ]. Measured Iosses in the self-sufficient mode at a nominal current of 1050 A required design values of about 1.8 W per pole, i.e., 1.7 mWI A-pole. 166 W. Barth, N. Fessler-Wilhelmi, W. Lehmann, and P. Turowski ,A \ •r des1gn volue w•th ~ IV I> .voo; / ~y.~Q "'<:' ~ I "' V!_)-___j__-------1 test volue fwtthoul i ony quench) I i 0 3 5 I 6 7 B, T Fig. 2. Characteristic values of conductor and compensating solenoids as air coils. Voltage tests have shown that these current leads can be operated up to a voltage of about 1.3 kV above ground. However, in the magnet tests a breakdown to ground occurred at a voltage of only 180 V. In this incident the 1-MJ stored energy of the magnet burned out parts of the current leads while the helium plasma beam burned a hole into the cryostat wall. The magnet coil itself showed no damage. This accident was obviously caused by a small undetected helium gas leak at the top of the current leads. The leak allowed the small space between the top and electrical ground to fall with helium gas. Since helium gas has a low breakdown voltage, the accident was inevitable. To avoid such accidents, metallic structural parts have been replaced by insulating glass fiber elements and additional selenium surge absorbers have been installed across the current leads as overvoltage protection. PERFORMANCE TEST OF THE SOLENOIDS At PETRA the two compensating solenoids will be operated in movable iron yoke doors. Before delivery and final mounting the two coils were pretested in a horizontal test cryostat as air coils at Karlsruhe. The characteristics of the solenoids under test conditions are shown in Fig. 2. Both magnets have been operated without any quench, i.e., without any training, up to an operating current limit of 1150 A with Superconducting Compensating Solenoids for the CELLO Detector Experiment 167 600 Fig. 3. Field decay and internal voltage characteristic of coils at 1150 A. Lower curve shows magnetic field decay from 3.8 T to zero; upper curve shows internal voltage curve of coil with a maximum value of 525 V. a central field of 3.7 T and a maximum field at the conductor of 4 T, i.e., 72% of the short sample value on the Ioad line of the coils. The discharge behavior of the magnet into an external resistor of 0.5 0, was studied at different currents. The data include field decay and valtage characteristics. Typical curves are shown in Fig. 3 for a maximum operating current of 1150 A . During discharge the coils partially reverted to the normal conducting state owing to the eddy current heating in the superconductor. About 70% of the stored energy was dissipated into the externalload. This is equivalent to a real temperature rise in the coil of only 30 K. The coils and the safety circuit were also testedunder real quench conditions. A Helmholtz coil pair was placed in such a manner that the windings of the magnet coil were located between this coil pair. A field pulse of about 200 T /s with a change in magnetic field of 0.3 T drove the magnet coil at 1150 A into a quench along the Ioad line. The result was an immediate appearance of coil resistance. The field decay initiated by the quench is similar in behavior to that noted after quick switchoff. These procedures have been repeated several times to test the reliability of the magnets. After these pretests the magnets were mounted in their permanent cryostats. The mounting arrangement of one of these magnets is shown in Fig. 4. Final tests of the magnets in the permanent assembly in Karlsruhe provided the same results as found in the pretests. CONDITIONS AND SPECIFICATIONS FOR CRYOGENIC SYSTEM The special character of the cryotechnical requirements lies in the restricted space available in the detector block and the limited access to the cryostats (see Fig. 1). These requirements for a very compact design of the cryostats and components resulted in helium lines, vacuum lines, and safety lines that are lang and complex. Figure 5 shows the basic arrangement of the helium cooling circuit; the cryogenic specifications are also indicated. The cooling concept selected provides foraparallel helium supply to the three superconducting magnets. 1t is necessary that the compensating coil cryostats can be moved axially while at the operating temperature. This requirement is satisfied by the installation of flexible line sections within the liquid helium supply lines between the "L-lines" and the "mainline" and within the 168 W. Bll'th, N. Fessler-WUhelmi, W. Lehmama, ud P. Turowski Fig. 4. Compensating solenoid during mounting into cryostat. warm return gas line (not indicated). This enables experimenters access to various measurement locations, such as spark chambers and liquid argon counters, without warming up the compensating coil system. CRYOGENIC DESIGN OF COMPENSATING COIL SYSTEMS The design of the cryostats for the compensating coils is shown in Fig. 6. The ring-shaped helium tank, with a liquid helium volume of approximately 25liters, was welded at the fronts and surrounded by an exit gas-cooled shield on all sides. The mechanical design has been basedonaxial forces of 15,000 N and radial forces of 5000 N; these specifications are predicated on a positioning accuracy for the coil of ±2 mm in the axial and ±0.5 mm in the radial direction relative to the iron yoke. Heat conduction to the liquid helium bath is kept small by use of special radial supporting elements between the cold shield and the helium tank. These consist of tubes and bolts arranged concentrically and in series to realize small cross-section areas and long distances for the heat ftow. The design is difficult because there is only approximately a 50-mm radial distance between the heliumtank and the cold shield. The heliumtank has been designed for a maximum pressure of 10 bar. Figure 7 shows the process ftow sheet for a compensating coil system. The helium is fed from the distribution box of the refrigerator through the transfer line system to the cryostat. The exit gas leaves the cryostat in four separate pathways, namely, through the two current Ieads (EL), through the cold shield of the inlet line, and through the cold shield of the cryostat. Part of the gas also passes in thermal contact with the safety line (SL). Alliinesare located within the same vacuum jacket (see section A-A, Fig. 6). The exit gas lines are heated to room temperature in separate ftows in a common electrical heater (690). After having passed through control valves and ftow measurement sections, the warmed return gas is fed to the compressor suction side of the refrigerator. The present design, i.e., heating all the return gas, allows individual cooling operation of the coils, shortens the overalllengths of the transfer lines, and obviates v -1 -·--·--·~ --- ·-- Compensating1__ . salenoi d A -- - · - · - Fig. 5. Basicarrangement of helium cooling circuit. Main solenoid 0.45 gls lor eurrent Iead 0.42 g LH•ts L-tine E - Lim it ol del ivery ------1 ,. KfK • -- • beam axis -·- : . -, CompensaFj ·Jng l~. Solenoid 8 2, 8 m M ... "' E 1 $ i· ä ii § l"" n i f ". I f f; ". g. ft 170 W. Barth, N. Fessler-WUhelmi, W. Lehmann, and P. Turowski Fig. 6. Compensating coil cryostat. the need for additional cold flexible line parts. The higher operational reliability of such a cryogenic system is rated against a lower overall efficiency of the system. Decreased efficiency results from a decreased use of the enthalpy of the cold return gas ftow from the compensating coil cryostats. The cryostat is protected against faults by means of a safety line integrated into the vacuum jacket of the L-line. Helium blow down is not permitted because of danger to the electronic equipment located within the detector block. In case of a pressure rise in a compensating coil cryostat the neighboring cryostats are protected by a nonreturn valve (NF 632) and by a corresponding valve control in the helium ftow to the cryostat (LCV 315) and a valve in the helium return ftow to the compressor of the refrigerator (FV 624). PERFORMANCE TESTS OF mE CRYOGENIC SYSTEMS Prior to delivery, cryostats and L-lines were tested together for tightness, function, and thermal Iosses at the manufacturer's site (Leybold-Heraeus/Köln). These tests were performed with a dummy cylinder welded in place of the solenoids and with liquid helium filling. After delivery the dummy helium tanks were replaced by the superconducting coils and the systems were completed with respect to the ftow Superconducting Compensating Solenoids for the CELLO Detector Experiment 171 Col dbox FV624 ; 9 Ef!J~ -X LCV 316 HCV 623 NF62!i to Comptessor L-----<- +--....J Air ' to second Compensating -- ---- --------, Solenoid ( 8 ) -- ---, I I I I' ' L----- - -------l - -------- -- ------- Fig. 7. Flow diagram of compensating coil systems. Legend: CSA, compensating coil A; CDL, cooldown line; SL, safety line ; TSL, transfer line (section J6-J8); EL, return gas cooled electricalleads; PS, phase separator; S, superconducting probe; Q , quenchdetector; and DR, discharge resistor. sheet. They were subjected to extensive cryotests at the Kernforschungszentrum Karlsruhe. The complete systems including the magnets (approximately 500 kg) were cooled down individually within 30 hr. The average coolant flow was 1 g/s at a maximum pressure of 1.2 bar abs with an enthalpy recovery of about 50% of the cold return gas. Since the complete finalliquid helium supply systemwas not available at either test location, it was only possible to measure the cryostat Iosses in a discontinuous mode via the helium evaporation rate while the supply was closed. Because of the restricted liquid helium volume (max. 16.5-liter reservoir capacity 171 W. Barth, N. Fessler-WUhelmi, W. Lehmann, and P. Turowski within the dome) there was only a limited period of stabilization. This resulted in higher loss values than expected for high liquid helium Ievels and a more pronounced dependence of the Iosses on the liquid helium Ievel (even at zero current) than can be explained by variations in the heat exchange surfaces and lengths of heat conduction, especially of the current Iead. Therefore, the results (see Fig. 8) of loss measurements performed on the compensating coil system B under different operating conditions represent maximum values. Under continuous ftow conditions at DESY, the dotted extensions of curves 3, 4, and 5 of Fig. 8 is the expected maximum Iosses at high Ievels. The increased evaporation Iosses which occur at 1050 A for liquid helium Ievels between approximately 30 and 45% inside the dome are a result of screw connections for the junctions between the normal conducting and superconducting parts of the current Ieads having a relatively !arge heat exchange area with the helium. The emersion of these two elements, located at different Ievels in the liquid helium bath, Ieads to a reduction in the evaporation of liquid helium. This effect becomes noticeable at the high current Ievel of 1050 A and a constant mass ftow through the lnsulal ion vacuum ä ,w s 10 ·6 mbar Temperature al lhe oullel of the cryostat shield 30- 60K 9 Specif icat ion value e design value 6 5 5 4 ---- - ...----".,----"'"-3 LHe- leve l ,'!. ~--~-+--+-~--------~~~~-80 70 nom ina'l Ievel 60 so 40 30 20 10 0 empt'y dome 1 - - - - - litt:O(J+) =110min - - - - - l Fig. 8. Evaporation Iosses of cryostat vs. liquid helium Ievel inside dome. Legend: Curve 1, · with dummy cylinder but without magnet and current Ieads; Curve 2, 0 with dummy cylinder and additional 5-W electrical heat Ioad inside liquid helium bath, simulating improved shield cooling; Curve 3, + with magnet and current Iead (I= 0, ni1e ad = 2 x 0.1 g/s) ; Curve 4, l;, with magnet and current Iead (I = 870 A (normal current), ni 1ead = 2 x 0.1 g/s) ; and Curve 5, 'l with magnet and current Iead (I= 1050 A (nominal current), 1e ad = 2 x 0.1 g/s). m Superconductlng Compensating Solenoids for the CELLO Detector Experiment 173 current leads of 2 x 0.1 g/s, the same value as for the nominal current of 870 A. In other words, the dashed line for curve 5 in Fig. 8 is approached with an increased rate of helium ftow through the current leads at the 1050-A level. In the continuous ftow mode at DESY, this increased ftow of heliumgaswill be available owing to the losses in the transfer line and the resulting supply of two-phase helium to the compensating coil system, as indicated in Fig. 5. The measured cryostat losses of the compensating coil system A were within approximately 10% of the losses for system B. As mentioned previously, a discharge of the magnet at 1050 A initiated a voltage ftashover at the current lead causing noticeable darnage which led to a breakdown of the insulation vacuum. Within a time period of 2 s, the pressure inside the helium tank and the insulation vacuum tank increased to the blowdown pressure of the firststage of the safety line (4 bar abs). This incident apparently caused a slight change in the insulating capability of the cryostat; the darnage could not be located and remedied with certainty without dismantling the welded components before delivery to DESY. Hence, after repair and reinstallation of the new current lead, the measured losses were higher by about 2 to 3 W for currents of zero and 1050 A. In spite of the higher losses of cryostat A, the sum of the losses (maximum values) stilllies within the values specified for the compensating coil systems. This is because of the lower values obtained for cryostat Band the two supply lines (L-lines) of 1.5 W each, including the loss of the couplings. Figure 9 shows the pressure rise inside the cryostats after discharge of the magnets. The curves indicate that for normal discharge and for flashover situations P abs • bar t [ l O S O A , Syst.A''' llOOA' i:- lOSOA " 870 A, Syst_ A 600A, Syst. B LH~ • 'Yolurnt' b4ol01t' sw•!cl\ ott 10 20 JJ '0 50 60 17 • 21 I t i me , s Fig. 9. Pressure rise inside cryostat and blowdown time after discharge of magnets. * Voltage ftashover incident in system A, destruction of current Iead inside vacuum tank, breakdown of insulation vacuum; **Voltage ftashover incident in system B, destruction of current Iead outside vacuum tank, breakdown of insulation vacuum; ***normal discharge of magnet with removal of 70% of the stored field energy and with the first safety valve closed. 174 W. Barth, N. Fessler-Wilhelmi, W. Lehmann, and P. Turowski encountered with the venting of the insulation vacuum by helium, the first blowdown stage of the safety line is sufficient protection for the system. The second relief valve to be opened at 6.5 bar abs, in addition to the first stage, can cope with the dynamic Ioads encountered when the insulation vacuum is vented with atmospheric air [5 ] and a simultaneous magnet quench without energy decoupling occurs. Functioning of the second valve was assured in one of the tests shown in Fig. 9. CONCLUSION The results of the pretests have shown that all requirements for the compensating coil systems can be fulfilled. The coils can be operated safely up to the design value without quench and the performance of the cryogenic system with regard to function, tightness, and Iosses matches the specifications. The compensating coil systems have been transferred to DESY, integrated within the CELLO detector block and the helium circuit. At the end of July 1979 the compensating coil systems were cooled down to 4.3 K for the first time in parallel with the main solenoid by the L'Air Liquide refrigerator in DESY. These tests confirmed that the overalllosses of the cryostats and transfer lines were within the specifications of Fig. 5 and that the increased loss noted previously within the compensating cryostat A was due to a contact between the heliumtank and the shield within the dome. This problern has now been eliminated. The initial Operation of the system at the installed CELLOexperiment has just started. Full rated current was attained without any problems. Further tests are in progress. ACKNOWLEDGMENTS The work reported is a joint eflort of mechanical, cryogenic, and electronic engineering groups of Kernforschungszentrum Karlsruhe, Institut für Technische Physik. The authors express their gratitude for the work done and cooperation experienced in the teams. The authors also thank the magnet group of CEN-Saclay/France for the fruitful collaboration, as weil as the assistance of Dr. Horlitz of DESY and Dr Oberlack of MPI München. REFERENCES 1. H. Desportes, in Proc. 6th Intern. Conference on Magnet Technology, ALFA, Bratislava, Czechoslovakia (1978), p. 474. 2. P. Turowski, M. Scherer, and W. Go II, in Proc. 5th Intern. Conference on Magnet Technology, Rome, Italy (1976), p. 541. 3. N. Fessler-Wilhelmi and K. P. Jüngst, "Ultrasonic Welded Multifilament Superconductors" Kernforschungszentrum Karlsruhe unpublished report. 4. J. M. Lock, Cryogenics 9(6):438 (1969). 5. W. Lebmann and G. Zahn, in Proc. 7th Intern. Cryogenic Engineering Conference, lPC Science & Technology Press, Guildford, England (1979), p. 569. D-2 CONSTRUCTION AND TEST OF TUE CELLO THIN-WALL SOLENOID H. Desportes, J. Le Bars, and G. Mayaux Centre d'Etudes Nucleaires de Saclay Saclay, France INTRODUCTION A large thin-wall superconducting solenoid has been constructed at Saclay and mounted on a large detector CELLO. This is one of the experiments installed on the e + e- colliding beam facility PETRA at DESY (Hamburg). The complete magnet system, in addition to this main solenoid, includes two compensating solenoids symmetrically located on each side of the main one, a thick 1000-ton iron shielding intended as a hadron filter, and a 300-W closed-cycle helium refrigerator supplying the three magnets. The two superconducting compensating coils have been designed and constructed at the ITP of Karlsruhe and are described elsewhere The major requirement for the main solenoidwas that it be very light weight or "transparent" to radiation. The amount of material allowed for the radial thickness of the complete magnet (including thermal shields and vacuum walls) was not permitted to exceed half a radiation length or 45 mm of aluminum. The design of the solenoid was prepared in close collaboration with the ITP in 1976. It was designed to Karlsruhe at the time of the CELLO physics proposal meet stringent requirements with the best available experience in superconducting magnet technology. The basic concepts of the design which were selected at that time These concepts were closely followed have already been reported elsewhere during the construction and have proved to be fully valid in view of the performance tests. The tests have demonstrated that such a design enables a minimum wall thickness to be achieved together with simple, reliable, and safe operational procedures. e]. eJ eJ. MAGNET DESIGN The two most significant points of the design are (1) the use of a high-purity stabilized aluminum, high-current conductor and {2) cooling by conduction from an external pipe in which a force fiow of two-phase helium is circulated. Stabilization of the conductor was considered essential to smooth out thermal effects of mechanical disturbances or of other instabilities and to provide sufficient heat capacity to the conductor so that it is intrinsically protected when a quench develops in the coil. High-purity aluminum is advantageous in the present application because its radiation length is six times higher than that of copper and its 175 176 H. Delportes, J. Le Ibn,_. G. May.u electrical and thermal conductivities are an order of magnitude higher than copper at low temperature. This last characteristic is particularly helpful for diflusing heat which may be produced locally. In case of a quench, it slows down the temperature rise due to the Joule eflect and speeds up the quench propagation through the entire magnet. Force-fiow cooling was preferred to bath cooling, mainly to reduce the amount of material involved in surrounding the coil with a helium container. Such a container would have required thick walls to withstand the large intemal pressures resulting from an aceidentat fiash vaporization such as could occur in the case of a vacuum break. Another advantage of force-fiow cooling is that it enables the magnet to be cooled down smoothly and uniformally without incurring thermal stresses due to large temperature gradients. Also, even though the cooling efficiency using exterior cooling may be sufficient for maintaining the magnet at the required temperature in normal operation, it becomes inadequate when large heat fiuxes are produced in the magnet; thus, it does not prevent quench propagation, which is essential for safety reasons. Other features of the design include a single continuous layer of conductors for the winding, a complete epoxy impregnation of all the elements of the coil package (bore tube, conductor, bandage, and cooling tube), the use ofthin aluminum-alloy shells for all the structural elements (bore tube, thermal shields, and vacuum tank), and an array of fiberglass struts distributed symmetrically at both ends of the solenoid as a support structure. MAGNET CONSTRUCfiON The first item needing special study was finding a conductor able to fulfill the above requirements. The conductor configuration selected is shown in Fig. 1. It consists of a reetangular 2.2 x 1.6-mm superconducting composite with the smallest possible ratio of Cu/Nb-Ti (1: 1), soft soldered onto a high-purity aluminum strip of 2.2 x 9 mm. The development and the entire fabrication of the aluminum stabilization were carried out at Saclay; the basic superconductor was ordered from Airco. Full details covering the fabrication and tests of this conductor have been reported elsewhere 4 ]. The total length of conductor used in the solenoid is 6700 m, which, in view of the limited lengths of the superconductor available, required four splices along the winding. Splices were made by cutring up part of the aluminum strip, overlapping the two ends of the composite over a length of 30 cm, and soft-soldering the three elements together in such a way that the cross section of the spliced partwas the same as that of the individual conductor; this method of fabrication will not disturb the winding regulari~. The electrical resistance of such a joint has been measured to be less than 2 x 10-9 n. A schematic cross-section of the magnet is shown in Fig. 1 where the various components are given with their respective dimensions. The actual winding procedure called for a carefully controlled process requiring special equipment. The winding line is shown in Fig. 2. The conductor was initially stored on a spirally wound spool (not visible in the photograph), with its axis in a vertical position. From this spool the conductor was driven through a tension device, a wrapping machine (for insulation), two epoxy wetting devices on each side of the wrapper, and was finally laid around the winding bore tube. The winding bore tube was driven by a horizontallathe and was fitted with a pressing device made of a ring clamped on the bore tube itseH. A series of jacks, assisted by means of sector pads, e· 177 Construdion and Test of the CELLO Thin-WaU Solenoid ,-----------------------------------------------Ou te r vac u um tan k , Ii II J ,--------------------------------------------------- 0 u ter sh iel d r------------------------------------------- Coo Iing tubes ,--------------------------- Mechanical bandage I Conductor Inner shield Inner vacuum tank -T-----------------------------------------------------+----------------------------------~--8 j; ~ n;r--[~~~~--""'DJ~ ~ L!'l ~ ~-------------------------------~20 N " (0 ' ~ ______________________________~ ~------------------------------------- 3560 --------------------------------------, ~---------------------------------------------- 4020 ---------------------------------------------~ Mechanical bandage Conductor _____________.....,._ Alum inum support Bore tube ---------·~>...l...lt-H..u..~ Fig. 1. (Top) Longitudinalcross section schematic of the CELLO solenoid and (bottom) detail of the coil package. Dimensions in millimeters. maintained a constant axial pressure on the edge-wound conductor as the winding progressed. This ensured a tight packing of the turns and prevented any tipping of the conductor. The conductor was insulated by double-wrapping it with silk tape of 80-#1-m thickness. (Silk was found mechanically stronger than glass when it was wetted and wrapped around the sharp cornered conductor.) The epoxy used was Ciba x 193/2594 C grade ; this epoxy remained liquid for weeks at room temperature. Progress of the winding is shown in Fig. 3 where the action of the pressing pads can be observed. 178 H. Desportes, J. Le Bars, and G. Mayaux Fig. 2. Winding line. After completion of the conductor layer, the surface was given extra insulation and an aluminum alloy stripwas wound over the insulation. This winding consisted of a continuous layer using the same winding line but without insulation and with the pressing device removed. At this stage, a thermally insulated shroud was mounted around the coil while it was still on the lathe. Curing of the epoxy, with infrared heaters placed inside the Fig. 3. Winding in progress. Construction and Test of the CELLO Thin-Wall Solenoid 179 Fig. 4. Finished coil with cooling pipe laid on its surface. shroud, was achieved at 150°C in 5 h; during this period the coil was continuously rotated on the lathe. After curing, the whole coil package was bonded tagether and formed a rigid structure. The next operation involved machining the inner bare of the bare tube to reduce its wall thickness to 3 mm from an initial thickness of 30 mm. The preformed cooling pipe was then glued and clamped on the surface of the solenoid using Stycast 2850 MT, a strong compound with good thermal conductivity. The pipe was constructed of a single piece of aluminum alloy tubing and laid along 16 straight runs, as shown in Fig. 4. This kept the length of the cooling loop to a minimum of 60 m resulting in a low pressure drop for the helium ftow, and also providing the same cooling path for each turn of the winding. The assembly of the coil in its cryostat required care, but the procedure was straightforward. The four cylindrical shells, consisting of the thermal shields and the vacuum tank, were positioned concentrically both inside and outside of the solenoid and were closed by end ftanges after the support system bad been securely tied to the ends of the solenoid. Multilayer insulation was applied to the entire surface between the shield and the vacuum tank. The complete assembly was carried out in a vertical position. The thermal shields are cooled with helium gas supplied from the refrigerator at 60 K. This cold gas is circulated in pipes and follows the same pattern as for the solenoid. The support system consists of two sets of fiberglass struts, one for centering and holding the weight of the coil and eventual asymmetric lateral forces, and the other for axial support. The arrangement of the struts is shown in Fig. 5. The thermal shield is suspended separately in a similar way. The last stage of the assembly required connecting all electrical and cryogenic circuits to the external transfer elements. This was done through a horizontal pipe (at one end of the solenoid) connected to a vertical chimney in which all the circuits were 180 H. Desportes, J. Le Bars, and G. Mayaux - -_·-:--:.: ===-=-=: _=;jt_=_=_ :_ =-=-=-~.-. lt ' 1/ /1 ·ft # i' /I end v i ew without cap_ back external vi ew with a cut i n the outer vacuum tank front Fig. 5. Arrangement of support struts. Fig. 6. Arrangement of equipment inside chimney. . _FRONT Construction and Test of the CELLO Thin-Wall Solenoid 181 Fig. 7. CELLO solenoid assembled inside the argon calorimeter and iron shield. collected according to the scheme discussed in the next section. The inside arrangement of the chimney is shown in Fig. 6. Figure 7 shows the completed magnet installed inside the large argon calorimeter used in the experiment at PETRA surrounded by the thick iron shield. CRYOGENIC CIRCUIT The three magnets installed in the experiment are supplied with cooling from the same refrigerator, but in different modes; the compensating solenoids are cooled in a pool boiling mode while the main solenoid is cooled in forced flow. The diagram of Fig. 8 shows only the cryogenic circuits for the main solenoid operation. A fraction of the helium flow is diverted at the last stage of the refrigerator and is subcooled and liquefied through a heat exchanger and a J-T valve. The two-phase flow is then circulated through the solenoid cooling pipe. From here it returns to an auxiliary reservoir situated in the chimney of the magnet. In the reservoir, part of the liquid is separated and part of the gas is extracted for cooling the current Ieads, which are connected to the terminals of the solenoids through leak-tight feedthroughs at the bottom of the reservoir. The remaining fraction of the two-phase flow is returned to the refrigerator cold box where the liquid is ultimately separated and mixed with the liquid directly produced by the second J-T valve from the main flow of the refrigerator. The reservoir in the refrigerator is used for transferring liquid to the compensating coils. 181 H. Desportes, J. Le Bars, and G. Mayaux . f!EFBIGEBATOR COLD WARM RETURN LIN...L_ LI NE 44 K r- · - - · PSV ... > ...w c PSV (I) L ., TA - Fig. 8. Cryogenic circuits for the solenoid Operation. Construction and Test of the CELLO Thin-Wall Solenoid 183 The entire system is operated by automatic control devices. The regulated parameters include the two-phase helium ftow in the solenoid cooling loop, both ftows of helium gas in the current Ieads, the liquid Ievels in the liquid helium reservoirs, the pressure of the cold helium return ftow, and the temperatures at the warm end of the current Ieads. The refrigerator itself, manufactured by Air Liquide, is a two-turbine type. 1t does not require liquid nitrogen and is also automatically controlled; accordingly, it requires a minimum of supervision and is designed for long periods of continuous Operation (-4000 hr). TESTS OF TUE SOLENOID Two series of tests have been performed on the completed solenoid, the first at Saclay without the iron shield and the second at PETRA with the entire system assembled for the experiment. The tests at Saclay were carried out without the refrigerator. Special cryogenic circuits were used which transferred helium from one dewar through the solenoid cooling loop and back to a second dewar. All the control and safety equipment was operated as in the final installation. The first test showed evidence of a fault in the solenoid. A permanent resistive spot was detected at all currents which caused the coil to quench repeatedly at a current of 1900 A (expected nominal current was 3400 A). The normal region had a resistance of 2 x 10-7 n, corresponding to about 10 cm of aluminum strip. The power dissipated in this resistance at the quench current was 0.7 W; this gives an upper Iimit to the heat which can be removed locally by the cooling scheme. After dismounting the solenoid, the defect was identified by X-ray techniques as an open break of the superconductor. Careful examination of the break revealed the presence of a metallump inclusion in the superconductor which clearly initiated the rupture. The break must have occurred during the winding but was not observed at that time since the two ends of the superconductor were still bonded to the aluminum strip and the entire conductor was wrapped with insulation. Since the broken conductor was located on the sixth turn from the end of the solenoid, the last six turns were removed from the winding and no other repair was attempted. This only shortened the solenoid by less than 0.5%. After reassembly, the magnetwas tested again and was charged to its maximum current. No training was experienced and no instability or jump could be observed during the current charging. Also, there was no temperature change nor mechanical deformation as noted by strain gauges measurements. The magnetwas quenched several times at a current of 3200 A, independent of the charging rate. This quench value is weil below the specified rated current of the conductor. In view of the above observations, the quench can only be explained by a degraded characteristic of the superconducting material in the area of the quench. The problern was located this time in the central region of the solenoid. This last assumption was further confirmed during the final test of fhe magnet installed on its experimental site at PETRA. This testwas carried out in July 1979. The two compensating solenoids, the huge iron shielding, and the refrigerator were also installed for the test. The complete system was cooled down from room temperature and filled with helium at 4.4 K in approximately 3 days. lt has been running continuously since then with remarkable stability and reliability. In this new environment the main solenoid could by charged up to 3000 A in less than 10 min, again with very stable behavior, but it was quenched at a current of 3025 A. This new 114 H. Desportes, J. Le Ibn, mtd G. Maya11X degradation can be attributed to two factors whose inßuence is only eflective when the current is limited by the H-1 characteristic of the superconductor. Theseare the magnetic field increase due to the iron and the operaring temperature of the solenoid. The first eflect is small, as the peak field is only increased by -5% in .the centrat region of the solenoid, but the temperature was effectively higher by as much as 0.2 to 0.3 K than during the test at Saclay. Such a temperature diflerence is responsible for a decrease in the current capacity of the superconductor by about 4 to 6%; this corresponds to the diflerence in the quench values. Further tests will confirm these conclusions. In view of these results, the practical field available for physics w~th CELLO is 1.3 T. An important result of the tests is the assurance gained on the complete safety of the magnet. Careful analysis of the quench behavior bad been carried out theoretically and was experimentally checked at three values of quench current during the tests. This analysis is presented elsewhere [5 ]. The results exceeded the most optimistic predictions; with respect to both longitudinal and transverse velocilies of quench propagation. During a quench at 3200 A, the normal resistance of the winding increased within a few seconds to a value four times higher than the external discharge resistor provided for protection and the temperature of the coil remained below 60 K; this shows that even without external protection the magnet would be intrinsically safe. lt is worth considering the role played by the conductor aluminum stabilization on the performance. Because of its very low resistivity, the diffusiontime constant for the current to transfer to the aluminum is large (of the order of 100 ms). Accordingly, at the front of the quench propagation all the current is in the small amount of copper and heats up the conductor rapidly, which, with the high thermal conductivity of the aluminum, helps to speed up the propagation. After this initial stage, the current transfers to the aluminum; the Joule effect is very much reduced and the temperature rise is slowed down. This permits a reasonably long time for the discharge of the magnet. CONCLUSION The concepts used in the design of the CELLO magnet have been justified. A better knowledge of the quench behavior has been gained. Technological experience on this rather new type of magnet construction will be useful in future designs. The operating current Iimitation experienced during the tests is a result of local degradation of the superconductor properties. The reason for this eflect is not known but it raises again the crucial need for complete quality control in superconducting material production. RE FE RENCES 1. W. Barth, N. Fessler-Wilhelmi, W. Lehman, and P. Turowski, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 163. 2. "Proposal for a 411" Magnetic Detector for PETRA," prepared by Desy, Karlsruhe, München, Orsay, Paris, and Saclay (July 1976). 3. H. Desportes, in Proc. 6th Intern. Conference on Magnet Technology, ALFA, Bratislava Czechoslovakia (1977), p. 474. 4. P. Genevey and J. Le Bars, in Proc. 6th Intern. Conference on Magnet Technology, ALFA Bratislava Czechoslovakia (1977), p. 1093. 5. W. V. Hassenzahl, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 185. D-3 QUENCHES IN THE SUPERCONDUCTING MAGNET CELLO W. V. Hassenzahl Los Alamos Scientific Laboratory, University of California Los Alamos, New Mexico INTRODUCTION The superconducting magnetCELLOwas tested with currents up to 3200 A at Saclay and has been installed at DESY in Harnburg where it will be used for particle physics experiments requiring colliding beams of electrons and positrons. The testing of this unique, large, one-layer solenoid provides an excellent opportunity to evaluate the theory of quench propagation under adiabatic conditions, that is, in a coil in which the conductors are not in direct contact with helium. In an early test of this coil, quenches occurred as a result of a broken conductor. This report describes the quenchesthat occurred, gives the details of the damaged conductor, and includes an analysis of the quenches. Observed axial quench velocities are compared to the calculated values based on both measurements and calculations of the thermal conductivity of the fabricated coil. The coil and conductor dimensions and characteristics are given in Table I. The coil and conductor are shown in Fig. 1, and a complete description can be found elsewhere CJ. Table I. CELLO Coil and Conductor Characteristics Interna) diameter, mm External diameter, mm Length, mm Thickness, mm Number of turns Design current, A Guaranteed current, A Operating current, A Short-sample critical current, A Protection resistance, 0 Inductance, H Cu to NbTi ratio Cu+ NbTi section, mm 2 High-purity Al section, mm 2 Residual-resistivity ratio of Al Thermal conductivity of high-purity Al at 4.2 K, W /cm-K 185 1656 1705 3414 24.5 1276 3400 2800 3100 -4000 0.16 1.06 1: 1 1.6 X 2.22 2.24 X 9 2=700 =60 186 W. V. Hassenzahl COOUNG TUBE MANOREL- WINOING TUBE Fig. 1. Layout and detailed cross section of the superconducting magnet CELLO. QUENCHES IN CELLO COIL Theoretical Background Superconducting coils are known to undergo transitions from the superconducting to the normal state. These transitions or quenches, which may begin at one or more places in the coil, propagate both along and perpendicular to the conductor. In the single-layer CELLO coil there is only one effective direction in which the quench can propagate perpendicular to the conductor. Quench Velocity Along a Conductor. The propagation velocity of quenches has been studied extensively [2-6]. The calculation by Broom and Rhoderick [2 ] gives the velocity of propagation, Vc, as a function of various conductor parameters and of the maximum temperature in the normal region by Ve = ) kp f(Om-20c)( OeOm(Om - Oe) C 112 J(k) =- - C Oe 1/ 2 G(O) (1) where C is the specific heat, k is the thermal conductivity, p is the resistivity, I is the current, Oe is the difference between the critical temperature and the operating temperature, and Om is the difference between the maximum temperature in the normal region and the operating temperature. Stekly and Hoag made the approximation that G(O) = 1, which is valid only after the centrat temperature of the normal region of a Nb Ti conductor reaches about 25 or 30 K [6 ' 7 ]. The velocities calculated with the actual conductor characteristics and G(O) = 1 are the dashed curves in Fig. 2. For comparison, the results of several measurements by Scherer and Turowski [8 ] are the solid curves in Fig. 2. The calculated velocities are typically higher than those measured. As the length of the conductor eJ Quenches in the Superconducting Magnet CELLO 60 I THEORY - - - EXPERIMENT--8 • 2T ....."' I I GORREGTION FOR SHORT SAMPLE E u g 40 / B~tT w / > I z f/ // 0 \ 1<[ (!) <[ a.. a: I / OBSERVED VELOCITY IN THE COIL I- 0 I PROPAGATION~~ ~ 187 20 / // / / a.. COMPOSITE ONLY I u z w / :::> 0 0 0 1000 2000 3000 4000 CURRENT,A Fig. 2. Ouench-propagation velocities for the CELLO superconductor-copper composite and the complete conductor with high-purity aluminum. Both theoretical and experimental values are shown. used by Scherer and Turowski is weil known, one can calculate the evolution of temperature within the sample and thus the propagation velocity as a function of time. The corrected velocity, which agrees better with the experiments, is marked with an asterisk in Fig. 2. This simple example demonstrates the need for a complete analysis of measurements made on small samples under laboratory conditions before applying the results to the performance of large coils. Quench Velocity Perpendicular to a Conductor. A quench propagatesnot only along but also perpendicular to a conductor. The perpendicular velocity depends strongly on the transverse thermal conductivity k l.. Wilson [9 ] shows that v l. = vc(kl.jkc) 112 • The question is: What is the magnitude of k1.? Of course, one can calculate k 1. and hope that the result is acceptable. One is reasonably confident in this calculation for a coil such as CELLO, but not for coils with round or cabled conductors nor for those cooled directly by helium. Table II gives the value of the axial thermal conductivity, kaxiah calculated from the known thermal conductivities of epoxy fiberglass and aluminum at 4.2 K. The thermal conductivities of copper and NbTi are also given in Table II, even though they contribute little to kaxiai because they are very small relative to that of aluminum. The averagethermal conductivity for a layered structure can be calculated from (2) where Ii is the length of a section of material, ki is the thermal conductivity of that section, and Rii is the contact resistance between materials. For CELLO, one can use 188 W. V. lbaeDDhl Table D. 'lbennal Conductivity of Materials in the CELLO CoU at 4 K Thermal conductivity, W/cm-K Material or configuration =60 =0.06 -0.05 1 X 10- 3 1.3 X 10- 3 1.0 X 10- 3 1.1 X 10- 3 Pure aluminum* Structural aluminum* Coppert:J: Superconductor:J: Epoxy fiberglass insulation:J: Conductor insulation* As-wound axial conductivity* * Measured value. t Calculated value. :J: Literature value. the values in Table II and a contact resistance, R, of approximately 5 cm 2 K/W C0 ] to calculate 2.24 0.278 kaxial=2.518 ( 60+1. 3 Xl0 3 +2x5 )-l =O.OllW/cm-K (3) To determine if this calculation was correct for the CELLO coil, a sample of ten insulated conductors was assembled in ablockthat simulated the actual coil. Table II gives the measured thermal conductivity of this sample. It is believed that the order of magnitude difference between the measured and calculated values of kaxiah is quite large because the test sample was not under compression; consequently the calculated value is closer to the characteristic of the assembled coil. One possible reason for this difference is the very poor cohesion between the epoxy and the thin layer of solder that forms the conductor surface. Quench Volume. The quench volume in two- and three-dimensional coils, Vz and V3 , respectively, can be characterized by (4) and (5) where t is the time, k" and ky are the thermal conductivities in the x and y directions, v" and Vy are the velocities in the x and y directions, and h is the thickness of a two-dimensional coil [ 11 ]. These equations are valid only until the quench reaches a boundary of the coil. In CELLO, for example, if an entire turn is normal before the quench reaches one end of the coil, the quench front will then propa~te only along the axis of the coil, and its speedwill be v" = vc(kd kc) 112, i.e., about 1lx, as fast as in the conductor. Here the normal volume increases linearly with time [dV/ dt oc Vc(kd kc) 112] until one end of the coil is reached. 189 Qoenches in the Sopercondoctiog Magnet CELLO Fig. 3. Voltages observed across the CELLO coil during a quench at 1900 A. Another possibility isthat at some time t 1 the quench reaches one end of the coil before a complete turn is normal. If this occurs, the volume is V= k 22(X) kc Vcf 1/ 2 2 + Vcf' f1 k ) (X kc 1/ 2 (6) ' until the quench completely encircles the coil. Quenches in the CELLO Coil Figure 3 shows the resistive voltages observed across the CELLO coil. These voltages indicate that the quenches started near one end of the coil. Of the three regions apparent in this figure, the first two appear to be quadratic, but the third is almost linear. These three regions are characterized in Table III. The linear slope during region 3, after t = 0.2 s, indicates that the quench is propagating in only one direction. The very low values of axial thermal conductivity in the coil and the slow increase in resistance indicate that this propagation is along the axis of the coil. As the quench was known to start near one end of the coil, the velocity can be calculated by assuming the quench front is moving axially along the coil in one direction. The velocity is given by il.R sA v =---=0.34m/s (7) ilt 27Tpr where il.R is the change in resistivity during the time ilt, s is the axiallength of a turn, A is the area of the aluminum, p is the resistivity of the aluminum at 4 K, and r is the radius of the coil. The slope during region 3 appears to deviate a bit from linearity in the last 0.3 s. This fact indicates that the maximum temperature is probably less than about 20 K, which is consistent with direct temperature measurements taken during the test and with computations made using the program OUENCH. Table 111. Voltages Observed Across the CELLO Coil During a Quench Quench region End time,* s Change in voltage, V 1 2 3 0.035 0.105 1.1 0.9 0.85 2.75 Change in resistance, n 4.9 4.4 1.5 X 10~ 4 X 10~ 4 X 10~ 3 Voltageltime characteristic Quadratic Quadratic Linear * Because the voltage changes very slowly at the beginning of a quench it is difficult to estimate exactly when it starts. Thus, the times may be low by as much as 0.007 s. 190 W. V. Hassenzahl From tbe calculated axial velocity of 0.34 m/s, tbe quencb sbould bave traveled about 1.2 cm or five turns during tbe first 0.035 s in region 1. The fact tbat tbe quencb actually started in tbe sixtb turn, as discussed later, supports tbis calculation and gives an axial velocity of about 0.40 m/s. Thus, at tbe end of -0.035 s, tbe quencb bad reacbed one end of tbe coil. For region 2, tbe quencb propagates along tbe axis of tbe coil in only one direction and continues propagating along tbe conductor, circumferentially around tbe coil. Finally, after 0.105 s, tbe quencb bad completely encircled tbe coil, traveting 2.64 m, to give a quencb velocity of 24 m/s along tbe conductor. This value, wbicb is plotted in Fig. 2, is considerably bigher tban eitber tbe calculated value or tbe value measured by Scherer and Turowski. As discussed above, tbe time at wbicb tbe quencb begins is not weH defined, but a cbange of 0.007 s will not reduce the velocity significantly. On tbe otber band, for the quench to meet itself on the opposite side of the coil it must transfer to the adjacent turns and then propagate along them. Thus the quench that begins in the broken section actually travels more than halfway around the coil before meeting itself in the fifth and seventh turns. Correcting for this effect would give a slightly higher velocity. The error bar on the point at 1900 A in Fig. 2 indicates these two effects. During subsequent tests of the CELLO coil a quench was initiated at 2550 A. It took 0.07 to 0.08 s for the quench to circle the coil, giving a propagation velocity of 33 to 38 m/s, which is also plotted in Fig. 2. If the above sequence were reversed, that is, if the quench bad encircled the coil after 0.035 s, then, instead of being quadratic, the voltage characteristic in region 2 would be linear. There remains a slight discrepancy between the axial quench velocity determined from the slope of the voltage in region 3 and the velocity found simply from the duration of the first region and the known position of the origin of the quench. This difference may be due to one or a combination of several factors. 1. The temperatures of the adjacent turns were elevated by Joule heating in the broken section before the quench began. This effect would reduce the heat required to achieve a given propagation velocity. Equivalently, (Jm and Oe of (1) would be reduced and Vc would be higher. 2. The actual duration of region 1 could have been 0.042 s, because the voltage increases slowly at the beginning of a quench. 3. The resistivity of the aluminum used to calculate the velocity may be too great. The third item deserves extra attention because an unrepresentative sample may have been used to measure the 4-K resistivity, which was found tobe about 4.5-5 nfi-cm. This value may be high by as much as a factor of 2 owing to excessive handling of tbe sample. The d V/ dt during the quench gives dR/ dt, which, in quench region 3, is proportional to the product pv. There is thus an ambiguity and neither p nor v can be found directly from region 3. However, as tbe velocity, 34 cm/s, is in error by at most 20%, the measured resistance can be off by at most 20%, neglecting magnetoresistance [ 12 ' 13 ], which is quite small in aluminum at fields below 1 T. The voltage of region 1 corresponds to almost 40 turns normal, even tbough it is certain that only 12 turns have portions that arenormal and that tbe total normal length must be between 10 and 12m. This difference is very puzzling. There is, however, a long delay between the time when tbe current transfers from the superconductor into the copper and when most of it bas transferred to the aluminum. 191 Quenches in the Superconducting Magnet CELLO I = 1900 A ORIGIN -1Lmm FROM ONE END DIFFU SION TIME FOR HIGH PURITY ALUMINIUM 1 =0.1 s --- ------ 2.0 >. --- IJJ ~ OBSERVED VOLTAGE 1.5 0 > 10 0.5 0 0 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 ,sec TI ME Fig. 4. A comparison of observed and calculated voltages across the CELLO coil during a quench originating in the broken conductor, six turns from one end of the coil. The measured resistances of the high-purity aluminum and copper are used. The only variable is the diffusion time for current into the high-purity aluminum. This process of current transfer is described by the well-known diffusion equation d 2J dx 2 = 1 dJ D dt (S) which gives an exponential solution for the penetration of current into a conducting slab. The characteristic time constant, T m • is given by 'Tm= 16L2 1097Tp (9) where L is the thickness of a slab conductor in cm and p is its resistance in !l-ern. The current penetrates 1 mm of aluminum having p = 5 n!l-cm in about 0.011 s. But 1 s is required for current to fully penetrate a 9-mm-thick aluminum slab. Though the diffusion of current into the aluminum conductor CELLO is more complicated than a simple exponential function, the program QUENCH was modified to give an exponential characteristic for the resistivity of the high-purity aluminum. The diffusion time Tm and the aluminum resistance were varied in the program. One of the closest approximations to the observed quench voltage is given in Fig. 4 along with an expanded version of the observed curve of Fig. 3. REPAIR OF CELLO COIL During these tests of CELLO a resistive section was observed at all currents. The voltage between turn 1059 and the end of the coil corresponds to a resistance of 2 J.'Ü or to a normal region about 12 cm long. 192 W. V. Hassenzahl Fig. 5. Enlargement of the in situ x-ray of the damaged section of the CELLO conductor. There is a gap of about 2 mm between the broken conductor ends. It was originally believed that this normal region was probably produced by overheating the conductor during the fabrication process when the copper superconductor composite was being soldered to the aluminum stabilizer. But it was possible that the conductor was broken so the coil was x-rayed. In CELLO, because there is only a thin layer of material that is relatively opaque to x-rays, =1.6 mm of copper superconductor composite, it is possible to detect a mechanical defect in the composite with this technique. Figure 5, a magnified copy of the original x-ray, shows a bad section of conductor, found in the sixth turn from one end of the coil. The broken conductor contained an inclusion of NbTi with a high oxide content. Inspection of the broken region clearly indicated that the break occurred before the epoxy in the coil was cured as there was some penetration of epoxy into the space between the composite and the aluminum near the break. Certainly the conductor was not broken before the soldering operation when the composite and the superconductor were joined. The last seven turns of the coil were removed and the void filled with an aluminum spacer and epoxy. FINAL TEST OF CELLO COIL AND CONCLUSIONS During the final test of CELLO a heater produced quenches at 2550 A and there were unprovoked quenches at 3200 A. The resistance of a section of conductor glued to the outside of the coil was used to monitor these quenches and to determine the final temperature after the quench. A temperature of 60 K was reached in less than 20 s after a 3200-A quench, indicating good thermal contact between all coil Quenches in the Superconducting Magnet CELLO 193 components. That the current is limited to 3200 A probably is because of a defect in the conductor similar to that which caused the break. This explanation is supported by the fact that the composite was ordered in a single length, =7000 m, but broke several times during fabrication, probably because of other inclusions. The analysis of this rather complex quench lends support to the use of existing theories to calculate quench propagation velocities and to predict the behavior of coils subjected to quenches. REFERENCES 1. H. Desportes, J. Leßars, and G. Mayaux, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 175. 2. R. F. Broom and E. H. Rhoderick, Brit. J. Appl. Phys. 2:292 (1960). 3. z. J. J. Stekly and E. Hoag, J. Appl. Phys. 34:1376 (1961). 4. W. H. Cherry and J. I. Gittleman, Solid-State Electron. 1:287 (1960). 5. L. Dresner, IEEE Trans. Magn. Mag-13(1):1328 (1979). 6. C. Meuris, "Protection d'une Bobine ou d'un Ensemble de Bobines Supraconductrices Lors d'un Passage a !'Etat Resistif," Saclay Rept. STIPE/76-80 (October 1976). 7. W. V. Hassenzahl, "Quench-Modifications and Documentation," Saclay Rept. SUPRA/78-61 (October 1978). 8. M. Schererand P. Turowski, Cryogenics 18:515 (1978). 9. M. N. Wilson, "Computer Simulation of the Quenching of a Superconducting Magnet," Rutherford High Energy Laboratory, Rept. RHEL/M 151 (1968). 10. M. Van de Voorde, "Results of Physical Tests on Polymers at Cryogenic Temperatures," CERN Rept. ISR-MA/75-38 (1975). 11. P. H. Eberhard, M. Alston-Garnjost, M. A. Green, P. Lecomte, R. G. Smits, J. D. Taylor, and V. Vuillemin, "Quenches in Large Superconducting Magnets," presented at 6th Intern. Conf. on Magnet Technology, Bratislava, Czechoslovakia, 1977. 12. F. Fickett, Phys. Rev. 83:1941 (1971). 13. V. I. Gostishchev and A. A. Drozd, Phys. Met. Metallurgy 39(6):168 (1975). D--4 CONSTRUCDON AND TESTING OF THE TWO-METER-DIAMETER TPC THIN SUPERCONDUCTING SOLENOID M. A. Green, P. H. Eberhard, R. R. Ross, and J. D. Taylor Lawrence Berkeley Laboratory Berkeley, Califomia INTRODUCTION The TPC experiment at PEP is one of several colliding beam experiments which use thin superconducting solenoid magnets. The CELLO experiment at PETRA in Hamburg, Germany and a detector for the ISR at CERN in Geneva, Switzerland use superconductors with an aluminum matrix 3 ]. The TPC magnet and a magnettobe used at Cornell University use the concept of a shorted secondary circuit to protect a high-current-density copper-based superconductor (4 ]. The Lawrence Berkeley Labaratory (LBL) has developed and tested the concept of proteering a large high-current-density solenoid magnet with shorted secondary windings [5 ]. This development work has been reported earlier (6 ]. The construction and testing of three test coils, two of which have a diameter of 1 m and one which has a diameter of 2 m have led to the construction of the TPC detector magnet. The TPC magnet is described elsewhere 8 ] as weil as in this report. This report presents the basic parameters of the TPC magnet, describes the steps of construction (and presents the results of the first tests of some of the magnet system subassemblies). c- e' TPC MAGNET BASIC PRINCIPLES The TPC magnet is different from almost all other large superconducting magnets which have been built to date. The data gathered on photons by the TPC detector will be enhanced by a magnet which is as transparent as possible to photons. As a result, the thin solenoid design concept evolved. The features which make the TPC magriet unique among large superconducting magnets are as follows: 1. The superconductor operates at matrix current densities which are much larger than conventionallarge superconducting magnets. 2. The TPC magnet is designed so that its quench is protected by a system of shorted secondary windings. These windings control the quench process and prevent bot spot formation. 194 195 Construction and Testlog of the TPC Thin Superconducting Solenoid lnsu lotor Squoshed 3 / 4" 0 0 Alum inum Tub ing 1100 Aluminum 0 I 10 I I 20 I I 30 I I Mill imeiers Fig. 1. Cross-sectional view of TPC superconducting coil package. 3. The TPC magnet is cooled by forced two-phase helium cooling instead of helium bath cooling. 4. The TPC magnet has an integrated coil and inner cryostat which is cast in epoxy resin (see Fig. 1). The TPC magnet has a warm inside diameter of 2.04 m. The outside diameter is 2.36 m in the center and 2.44 m at the ends. The length of the cryostat, which encloses the coil, electricalleads, and coil cryogenic system, is 3.84 m. Access to the coil, the electrical services, liquid helium, signal wires, and vacuum is restricted to a small portion of the outside corner of one end of the cryostat vacuum vessel. The TPC magnet refrigeration system (refrigerator cold box and control dewar) is connected to the magnet by about 18m of transfer line running under a wall of radiation shielding. The magnet is monitored by a small computer. Table I shows the basic parameters of the TPC magnet as it operates at design current in its iron return yoke. When the TPC magnet is operated in its iron yoke, it behaves like an infinite solenoid which has a peak induction in the coil which is scarcely different from the uniform induction within the magnet bore. The TPC magnet is designed to produce a field near the center, uniform to better than 1 part in 1000. High field uniformity is required in order that the time projection chamber detector functions properly. The design current density, J, in the superconductor is ver~ high for a magnet which operates at a stored energy, E, of 11 MJ. As a result, EJ product is over fifty times that of conventional large superconducting magnets [9]. Table I. Basic Parameters of TPC Magnet in Its Iron Return Yoke Coil diameter, m Coillength, m Number of turns Design central induction, T Design current, A Magnet coil inductance, H Current density J in superconductor matrix at design current, A/m 2 Magnetic energy stored at design current, J EJ2 productat design current, JA 2fm 4 Maximum charge voltage, V Minimum charge time to design current, s 2.172 3.295 1772 1.5 2225 4.41 6.87 X 108 10.91 X 106 5.15 X 10 24 5 2000 196 M. A. Green, P. H. Eberhard, R. R. Ross, and J. D. Taylor Fig. 2. Winding of superconducting coil onto epoxy-cast ultrapure aluminum layer. CONSTRUCTION OF THE TPC MAGNET Construction of the TPC magnet began during the summer of 1978 with the fabrication of a 9.55-mm-thick 1100-0 aluminum bore tube, which is one of the shorted secondary circuits used for quench protection (see Fig. 1). A layer of 600 turns of ultrapure aluminum, insulated turn to turn with dacron cord, was wound on the bore tube. The ultrapure aluminum, which is in the form of a 3-mm-diameter wire, has a residual resistance ratio at 4.2 K of at least 1500. A great deal of care bad to be used while winding the ultrapure aluminum because this material deforms at almost no stress, resulting in increased resistivity of the aluminum at low temperature. Once the ultrapure aluminum layer was wound, it was vacuum impregnated in epoxy resin. The cast ultrapure aluminum layer provided a firm platform on which to wind the 1.0 x 3.7-mm Formvar insulated superconductor. The superconductor, which has a copper to superconductor ratio of 1.8, has over 2000 25-t.tm-diameter Nb-Ti filaments, which are twisted one twist every 50 mm. This conductor will carry at least 3200 A at 4.2 K and 2.0 T. Figure 2 shows the superconductor being wound onto the cast ultrapure aluminum layer. The 1772 turns of superconductor was wound in two layers under tension of 690 N (155 1b). This prestrains the superconductor to allow it to match the thermal contraction of the surrounding aluminum. The superconducting coil has an electrical center tap between the two layers. (See Fig. 1 for the location of the superconductor.) A layer of ftattened 19.1-mm 00, 3002 aluminum tube with 2-mm-thick walls is wound over superconducting coil. The aluminum tube serves a dual function. One-third of the turns carry two-phase liquid helium to cool the magnet, and the layer of aluminum tube serves as an elastic support for the coil system. The latter function Iimits the total strain of the coil package when magnetic forces are applied. The superconductor and ftattened cooling tube were vacuum cast with epoxy resin to form a rigid integrated structure. CONSTRUCTION OF THE CRYOSTAT The TPC magnet aluminum cryostat vacuum vessel has an ID of 2.04 m and an OD at the ends of 2.44 m. The centrat section of the outer cylinder is recessed for some drift chambers which are about 3 m long. The overalllength of the cryostat is 3.84 m. The TPC cryostat vacuum vessel not only provides vacuum for the superin- Construction and Testing of the TPC Thin Superconducting Solenoid 197 Support Pod Outer Vocuum Vessel Superconducting Coil Pockoge T PC Pressure Vessel Fig. 3. Cross-sectional view of coil package within its cryostat (note support rod). sulation but the inner wall is also an 11-atm pressure vessel which holds the argon-methane mixture needed for the time projection chamber. A cross section showing the coil package within the cryostat is shown in Fig. 3. This figure also shows one of the compression support rods. The TPC magnet coil package is insulated with multilayer insulation and has a liquid nitrogen shield. The insulation and the shield are not shown in Fig. 3. The superconducting coil package is connected to the outer cryostat vessel with a bicycle-type support system of 25-mm-diameter fiberglass epoxy rods. The support systemwill cary a Ioad of 2 x 105 N (20 tons) in any direction. The cold mass of the coil package is about 1500 kg. The total refrigeration needed to cool the TPC magnet coil package is estimated tobe between 10 and 15 W at 4.4 K (not including the gas used to cool the 2500-A electricalleads). CONSTRUCTION AND TESTING OF THE TPC CRYOGENIC SYSTEM The TPC magnet is cooled with two-phase helium carried in the tubes around the coil (see Fig. 1). The principle behind the cryogenic system is described in detail elsewhere C0 ] . The TPC magnet cryogenic system consists of a 200-liter control dewar, a conditioner dewar, and 85 m of liquid nitrogen shielded transfer lines which carry both liquid nitrogen and two-phase helium to the TPC magnet and two 1.8-m-long compensating solenoids. The control dewar contains the helium pump described by Green et al. [10] and a copper tube heat exchanger. The control dewar reduces the inlet quality of the two-phase heliumentering the cryogenic distribution system and the control dewar is a controlled buffer volume of liquid helium between the refrigerator and the Ioad. The TPC magnet control dewar with its heliumpump is shown in Fig. 4. The control dewar, helium pump, and the primary transfer lines were tested before final connection to the magnet. The measured control dewar heat leak varied from 2 to about 6 W depending on the liquid Ievel in the control dewar. Initialtests of the transfer line performance showed a heat leak of about 0.3 to 0.4 W /m. This heat leak is high and there are expectations to reduce it through the elimination of 198 M. A. Green, P. H. Eberhard, R. R. Ross, and J, D. Taylor Fig. 4. TPC magnet control dewar and helium pump. oscillations and improved insulation. The helium pump was operated at mass ftow rates from 8 to 40 g/s circulating through the control dewar and transfer line system. Helium pump adiabatic efficiencies of about 50% were measured. MAGNET POWER SUPPLY, QUENCH PROTECTION SYSTEM AND ELECTRONIC DATA LOGGERS The TPC magnet power supply can provide the TPC magnet with up to 3000 A at 10 V. The power supply can charge the TPC magnetto full design field in about 40 min. The TPC power supply is a six-phase supply which is regulated on the primary side of the transformer. The power supply has been successfully tested at full current. The quench protection system consists of four distinct elements: (1) the 9.5-mm-thick 1100 aluminum bore tube, (2) the layer of ultrapure aluminum, (3) a SCR "circuit breaker" with a varistor across the Ieads, and (4) a capacitor bank which discharges into the center tap between the two layers of the coil. The first two quench protection systems are passive, the latter two elements are dynamic quench protection systems which require the quench tobe detected quickly. Bothof the dynamic quench protection methods have been tested on a 2-MJ thin solenoid [ 12 ]. A microprocessor data Iogger is used to monitor twelve channels of data simultaneously every millisecond. The data Iogger is programmed to take data in a Construction and Testlog of tbe TPC Tbin Superconducting Solenoid 199 set sequence for 10 s after a quench has been detected. The magnet coil package has the following instrumentation built into it: (1) ten small coils, which can be used to initiate a quench and measure the velocity of normal region propagation, (2) five silicon diode temperature sensors, (3) five coils closely coupled to the superconducting coil which measure the change of magnetic flux during a quench, and (4) three voltage taps at the coil ends and at the center tap. The data from the instrumentation built into the coil package is transmitted to the data logger. The data stored in the data logger are then processed by the TPC experiment computer. CONCLUSIONS Tests on the completed TPC magnetwerein preparation in mid-1979 and the extent of passive quench protection afforded by the design was not known. Tests of the TPC magnet without iron are scheduled during the fall of 1979. Tests of the TPC magnet with iron at SLAC are scheduled to occur in the spring of 1980. ACKNOWLEDGMENTS The authors wish to acknowledge the efforts of A. Barone, D. Coyle, and W. Wenzel of the LBL assembly shop. They wish to acknowledge the engineering effort of P. Miller, W. Burns, and B.Garfinkel of the LBL Mechanical Engineering Department. They thank C. Covey and H. Van Slyke for continuous effort expended building the magnet cryogenic distribution system. G. Gibson and R. Smits are acknowledged for their rote in the development of the TPC Magnet electronics systems. This work was performed under auspices of the United States Department of Energy. REFERENCES 1. "Proposal for a 4'71" Magnetic Detector for PETRA (CELLO)," DESY, IEKP Karlsruhe, MPI München, Orsay, Paris University, Saclay (August 1976). 2. M. Morpurgo, Cryogenics 17(2):89 (1977). 3. M. Morpurgo and G. Posso, Cryogenics 17(2):87 (1977). 4. M. A. Green, in Proc. 6th Intern. Conference on Magnet Technology, ALFA, Bratislava, Czechoslovakia (1978), p. 429. 5. R. C. Wolgast, H. P. Hernandez, P. R. Aron, H. C. Hitchcock, and K. A. Solomon, in Advances in Cryogenic Engineering, Vol. 8, Plenum Press, New York (1963), p. 601. 6. M. A. Green, Doctoral Dissertation, University of California, Berkeley, California (1977). 7. M. A. Green, W. A. Burns, P. H. Eberhard, G. M. Gibson, P. B. Miller, R. R. Ross,R. G. Smits,andJ. D. Taylor, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1979), p. 86. 8. M.A.Green,P.H.Eberhard,J.D. Taylor, W.A. Burns,B. Garfinkei,G.H. Gibson,P. B. Miller,R. R. Ross, R. G. Smits, and H. W. Van Slyke, IEEE Trans. Magn. Mag-15(1):128 (1979). 9. P. H. Eberhard, M. Alston-Garnjost, M. A. Green, P. Lecomte, R. G. Smits, J. D. Taylor, and V. Vuillemin, in Proc. 6th Intern. Conference on Magnet Technology, ALFA Bratislava, Czechoslovakia (1978), p. 654. 10. M. A. Green, W. A. Burns and J. D. Taylor, in Advances in Cryogenic Engineering Vol. 25, Plenum Press, New York (1980), p. 420. 11. J. D. Taylor, M. Alston-Garnjost, P. H. Gibson, M. A. Green, B. Pardoe, M. Pripstein, R. R. Ross, and R. G. Smits, IEEE Trans. Magn. Mag-15(1): 855 (1979). D-5 SUPERCONDUCI1 NG MAGNET SYSTEM FOR TUE SPIRIT COSMIC RAY SPACE TELESCOPE M. A. Green and J. M. DeOiivares Lawrence Berkeley Laboratory Berkeley, Califomia and G. Tarle, P. 8. Price, and E. K. Shirk University of Califomia Berkeley, Califomia INTRODUCfiON The identity of the source of the cosmic radiation is one of the oldest and most interesting unanswered questions of 20th-century physics. While it has become increasingly clear that these energetic particles owe their existence to some of the most violent processes that occur in our galaxy (e.g., supemovae), a detailed understanding of the cosmic ray source and the conditions of galactic propagation has not been achieved. Of particular interest in this regard is the isotopic composition of the cosmic radiation since nuclear abundance anomalies would provide the most exciting clues as to their nuclear origin. The iron isotopes provide the most fruitful candidates for such a study since they are both abundant and are least modified by galactic transport. To date the most accurate isotopic studies of the iron group cosmic rays CJ have ruled out large deviations from solar system source composition. In order to achieve a convincing separation of the isotopes of iron it is necessary to design an instrument which can collect over 104 iron nuclei and achieve a mass resolution of u :5 0.15 amu. Three important developments have made the design of such an instrument possible: (1) With considerable impetus from the new generation of accelerators, magnet and cryogenic technology has reached the stage where very large volumes can be filled with uniform fields in excess of 2 T; (2) A track recordingplasticdetector made of CR-39 has been developed that is sensitive to minimum-ionizing particles of the so-called very heavy group (20 :S Z :S 30) has very good charge resolution, yields etched tracks of very high optical quality, and can be made in films thin enough that multiple Coulomb scattering can be negJ.ected; and (3) the advent of the space shuttle will make it possible to lift payloads weighing many tons into orbit for periods of several weeks. eJ 200 Supercoadudin1 Mapet System for tbe SPIRIT Coamic Ray Space Telescope 201 One of the authors (GT) has conceived of an instrument (SPIRIT) which capitalizes on these recent developments and can achieve the required collecting power and resolution in a 10-day shuttle ßight using only passive components. A three-tiered passive hodoscope consisting of track-recording plastic with a thin (60-"m) centrallayer will record the trajectories of cosmic ray particles through a magnetic field with an average strength of 2 T. These particles will be traced to their end of range in a stack of CR-39 where their charge will be determined by measurements of etched cone length. The measurement of magnetic rigidity in combination with the measurement of range will b"e used to determine particle mass. THE SUPERCONDUCfiNG MAGNET In order to achieve the resolution and collecting power necessary to meet the experimental objectives of SPIRIT, a superconducting magnet with an average field of 2 T over a control volume of 1 m 3 is needed. The control volume must be clearly accessible to cosmic radiation entering from a polar angle from 0 to about 45° (see Fig. 1). Spatial gradients need to be kept below 4 T/m, so that shifts in detector orientation expected during flight will not result in a degradation of resolution. The entire apparatus must be contained in the space shuttle orbiter cargo bay which has a dynamic envelope diameter of 4.57 m. The proposed magnet consists of six coils. The four inner coils generate a uniform high field over the control volume. The two outer coils will generate a smaller field over a larger volume so that the entire assembly will have a zero net dipole moment, which results in a negligible distant field. The design of the SPIRIT magnet system is similar in concept to the highcurrent-density type of magnet which has been under development for the last four years 5 ]. This development work has come to fruition with the construction of the 2-m-diameter, 3.3 m long, 1.5-T thin solenoid which is designed to operate at a currentdensityof7 x 108 A/m 2 andatastoredenergyofgreaterthan 10 MJ[5 ]. This design concept, which is used in the TPC solenoid [5], is particularly applicable for use in space. The SPIRITmagnet has the following characteristics: (1) Intrinsically stable high-current-density superconductor is used, (2) quench protection is based on the use of LBL shorted secondary concept, (3) cooling of the superconducting magnet is e- 1~ Fig. 1. Cross-sectional view of SPIRITexperiment showing the six-coil superconducting magnet system. 202 M. A. Green, J. M. DeOiivares, G. Tarle, P. B. Price, and E. K. Sbirk done with pumped two-phase helium, and (4) the magnet coil, the shorted secondary, and the cooling system are integrated into a single package which is contained within a single cryostat vacuum vessel. Large currents are required in the four inner coils in order to generate the 2-T induction over the volume of the experiment and meet the requirement for clearly accessible solid angle. As a result, peak inductions within the superconductor will approach 8 T. Therefore, it is proposed that the four inner coils be wound with prereacted multifilamentary Nb 3 Sn. The two outer coils, which carry lower currents would use multifilamentary Nb-Ti conductor. Electrically, the six coils are connected in series; the loop is closed by a persistent switch. The SPIRITmagnet would operate in the persistent mode for the entire 10 days of the shuttle ßight. Table I shows the basic parameters of the six coils. The placement of the coils in the magnet is shown in Fig. 1. There are three types of coils (there are two coils of each type). The four inner coils carry over 11 x 106 A in order to generate 2 Tover the experimental volume of 1 m 3. As a result, the use of Nb 3Sn conductor is proposed. If one reduces the average induction in the experimental region from 2 to 1.5 T, the current in the inner coils is reduced to 8.2 x 106 A. The peak induction in the coil is reduced to 6.0 T which permits Nb-Ti conductor to be used instead of Nb3Sn. Table II presents the electrical parameters of the six coil magnet system shown in Fig. 1. The proposed SPIRITmagnet has an inductance of 85.1 H. When the central induction is 2 T, the magnetic energy stored in the coil system is just over 34 MJ. It is proposed that the coil superconductor be operated at a current density of 3 x 108 A/ m 2 • The proposed conductor current density is about six tim es higher than the conductor current density normally used in a magnet with a 34-MJ stored energy. (Typical superconducting magnets with a stored energy of 34 MJ use cryostable conductors with a matrix current density of less than 5 x 107 A/m 2 .) The high conductor current density is necessary in order to reduce the mass of the magnet system. Table 111 presents the parameters for the superconductor proposed for use in the SPIRIT magnet. The proposed operating current density and high stored energy result in a high EJ 2 product, where E is stored energy and J is superconductor matrix current density. Thus, it is proposed that a well-coupled secondary circuit made from very pure aluminum be used for quench protection. The shorted secondary quench Table I. Basic Parameters of tbe Three Types of Coils Proposed for the SPIRIT Superconducting Magnet System Parameter Inside diameter, m Outside diameter, m Distance from center, m Coil package length, m Coil package total current, A Number of turns Current per turn, A Peak induction in the coil package, T Superconductor type Coils 1Aand lB* Coils 2A and 2B* Coils 3A and 3Bt 0.800 1.100 0.600 0.400 2.484 X 106 2760 900 -8.2 Nb 3Sn 1.100 1.400 0.750 0.250 3.096 X 106 3440 900 -7.2 Nb3 Sn 1.892 2.292 1.250 0.200 -1.620 X 106 1800 -900 -4.0 Nb-Ti * Inner coils which produce the 2-T field within the experimental volume. t Outer coils which cancel the dipole moment generated by the inner coils. Superconducting Magnet System for the SPIRIT Cosmic Ray Space Telescope 203 Table II. Electrical Parameters for the SPIRIT Superconducting Magnet System (Six Coils Hooked in Series) Integrated average induction within the experiment, T Experimental volume, m3 Magnet system design current, A Magnetsystem self-inductance, H Magnet system stored energy at its design current E, J Current density in the superconductor matrix J, A/m 2 El 2 productat the design current, JA 2 /m 4 2.0 -1.0 900 86.2 34.3 X 106 3 X 108 3.1 X 1024 protection system is used on the 2-m-diameter TPC solenoid which has an EJ 2 product of 5.4 x 10 24 JA 2 /m 4 • The shorted secondary concept affects the quench process in the following ways: 1. The shorted secondary causes the coil current to shift from the coil to the secondary circuit. As a result, there is less current in the coil to contribute to the conductor bot spot. 2. The shorted secondary circuit absorbs a substantial amount of the magnet stored energy. In the proposed magnet system, the shorted secondaries are expected to absorb about 70% of the magnetic energy (this energy will be shared by all of the coils, not just the one that went normal). 3. The shorted secondary will cause "quench back" in the other coils when one of the six coils turnsnormal through ordinary quench propagation. ("Quench back" is the process by which superconducting magnet coils are driven fully normal by the shortened secondary winding [4 ].) Quench back is a key element in the protection of thin high-current-density solenoids which have been built at LBL. The shorted secondary circuit would be insulated from the superconductor. It is desirable that inductive coupling between the coil and the secondary circuit be maximized. It is proposed that the shorted secondary circuit be made from ultrapure aluminum (0.99999 pure or better) which has a residual resistance ratio at 4.2 K and 0 T of about 2000. Aluminum has a much lower magnetoresistance at 8 T than Table III. Properties of Superconductor Proposed for SPIRIT Superconducting Magnet System Nb-Tit Uninsulated matrix dimensions, mm Insulated matrix dimensions, mm Insolation type Copper-to-superconductor ratio Bronze-to-unreacted-Nb ratio Number of filaments Filament diameter, m Twist pitch, mm Averagematrix resistivity at 4.2 K (!1-m) Current capacity at 4.2 K, A at ST at lOT 3X1 3.1 X 1.1 epoxy -1 to 1 -2.8 to 1 >30,000 -4 -so -10-9 3900 1100 * Multifilamentary Nb 3 Sn is proposed for the four inner coils 3X1 3.1 X 1.1 Formvar -1.8 X 1 NA -2000 -2S -2 -so X 10- 10 1600 1A, 1B, 2A, and 2B. This conductor has some copper in the matrix. t Multifilamentary Nb-Ti is proposed for the two outer coils 3A and 3B. M. A. Green, J. M. DeOUvares, G. Tarle, P. 8. Price, and E. K. Shirk 204 Aluminum shorted secondory circuil ·® Nb1 Sn supe rconduct ing coils Eleclri col insulotion • \ ® CJ 1\1\r ® ® \ Helium cooling tubes ® \ \ \ \ \ SC ALE 10 \ \ 15 \ 20 CM \ !.\\ Fig. 2. Cross section of proposed inner coils (coils 1 and 2) for the SPIRIT magnet. copper. lt also has one-third the density. At full field, one can expect the shorted secondary circuit to have a residual resistivity ratio greater than 300. As a result, the shorted secondary circuit is expected to have a time constant in excess of 30 s. If the coupling between the coil and the shorted secondary circuits meets expectations, effective shifting of the coil current will occur. The proposed coils will have the superconductor, shorted secondary, and a forced-ftow tubular cooling system combined into an integrated package. The proposed superconducting coils are designed to be weil insulated with insulation to ground adequate to 10 kV or more. Therefore, the superconductor, shorted secondary, and cooling system aretobe cast in epoxy resin. This technique has been used successfully in the TPC solenoid and three 1- and 2-m-diameter test coils [6 ] . Figure 2 shows a proposed arrangement of superconductor, shorted secondary circuit, cooling tube, and mechanical support inside of the two inner coils. The shield coils not only effectively eliminate magnet moment, but also greatly reduce stray magnetic field in the shuttle bay. lt is proposed that the superconducting magnet be located at the rear of the space shuttle bay. The expected stray magnetic induction in a region normally housing astronauts is expected to be around 10-4 T. 1t is proposed that coil 1 and 2 (in each half) be attached directly (see Fig. 2). A compressive force of 7. 7 x 105 N (77 metric tons) is expected between the two halves (between coils lA, 2A, 3A and coils lB, 2B, 3B). This forcewill be carried with the cold column struts between coil lA, 2A and coil lB and 2B. The columns are arranged so that there is full access of the cosmic rays to the experiment. A tensile force of 5.8 x 106 N (580 tons) is expected between each of the two outer coils and their companion inner coils. This force will be supported with a continuous web of metal between the outer and inner coils. The six coils are expected to act as a rigid frame which will have a cold mass of about 4000 kg (the helium tanks and the coil cryogenics will attach directly to this frame). THE CRYOSTAT AND CRYOGENIC COOLING SYSTEM The proposed superconducting magnet coils will be cooled using two-phase helium pumped through tubes in the coil package (see Fig. 3). The two-phase helium will be circulated from a separate helium storage tank located at the end of the experiment (see Fig. 1). Forced tubular cooling offers a number of advantages over the more conventional bath cooled systems CJ: 1. Tubular cooled systems can be cooled easily from room temperature by a refrigerator which is external to the cryostat system. Superconduding Magnet System for the SPIRIT Cosmic Ray Space Telescope 105 ----1------_:• •=·~ --' Gas cooled rod1at1on sh1elds r ·------ -----,I I I I I I I I I I I ____________ j Fig. 3. Simplified schematic diagram of SPIRIT magnet helium tank and distribution system. L ~~0 ~e; h~u: ~e~t :c;:n~r- "fc~o-:;a~ storage tank \ vocuum boundary I I --I 2. Only a small fraction of the liquid helium is in direct thermal contact with the superconductor at any one time during a quench. The tubular cooling system can contain the pressure rise owing to this small amount of helium. Helium boil-off during a quench is orderly and weil controlled. 3. In conventional systems, diamagnetic repulsion of the helium in a weightless environment would result in a loss of liquid cooling capacity [8 ]. Tubular cooled coils would contain two-phase helium which is in direct contact with the coil at all times. 4. The design of the cryostat is simplified. Many of the cryogenic safety problems found in conventional bath systems are eliminated. The forced two-phase tubular cooling system requires a helium pump to circulate the helium from the helium storage tank located at the end of the magnet. The entry to the pump is located near the zero field point in the tank. As a result, diamagnetic repulsion insures that liquid heliumwill always be delivered to the pump entry [9 ]. The pump proposed is similar to a reciprocating bellows-type pump which The pump will be designed to pump 1.0 g/s across a has been under development pressure rise of 2 x 104 Pa (0.2 bar). lt is expected that such a pump will require about 1 to 1.5 W of refrigeration. (This is equivalent to 1.4-2 liters/hr of liquid helium from the tank.) A 2500-liter helium tank can be built into the end of the coil as shown in Fig. 1. The coils and tank would be thermally isolated with fiberglass epoxy spacers. The coils and tank would be insulated with a combination of superinsulation and shields which use the helium boil-off from the tank. A total heat leak of 2 to 3 W is expected into the 4-K region producing a helium boil-off rate of about 5 to 6 liters/hr. The total helium inventory in the tank should be enough for about 20 days. n. CONCLUSIONS The proposed SPIRIT superconducting magnet system appears to be within the state of the art. A coil system which uses multifilamentary Nb 3 Sn and Nb-Ti is similar to the proposed fusion magnets. Forced two-phase cooling and quench protection using shorted secondary circuits has been demonstrated. Forced-cooled high-current-density superconducting coils are weil suited to space application. The development of this technology and the space shuttle make possible the study of some of the fundamental physics which occurs in deep space. 206 M. A. Green, J. M. DeOiivares, G. Tarle, P. 8. Price, and E. K. Sbirk ACKNOWLEDGMENTS The authors acknowledge work done by P. H. Eberhard and others of the Lawrence Berkeley Laboratory. Much of the research which has led to this report was performed under the auspices of the United States Department of Energy and the Space Seiences Laboratory of the University of California at Berkeley. REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. G. Tarle, S. P. Ahen, and B. G. Cartwright, Phys. Rev. Lett. 41:771 (1978). B. G. Cartwright, E. K. Shirk, and P. B. Price, Nucl. Instrum. Meth. 153:457 (1978). M. A. Green, in Advances in Cryogenic Engineering, Vol. 21, Plenum Press, New York (1975), p. 24. M. A. Green, Doctoral Dissertation, University of California, Berkeley, California (1977). M. A. Green, P. H. Eberhard, J. D. Taylor, W. A. Burns, B. Garfinkel, G. H. Gibson, P. B. Miller, R. R. Ross, R. G. Smits, and H. W. Van Slyke, IEEE Trans. Magn. Mag-15(1):128 (1979). M. A. Green, P. 8. Miller, and W. F. Wenzel, in Nonmetallic Materials and Composites at Low Temperatures, Plenum Press, New York (1979). M. A. Green, W. A. Burns, and J. D. Taylor, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 420. W. L. Pope, G. F. Smoot, L. H. Smith, and C. E. Taylor, in Advances in Cryogenic Engineering, Vol. 20, Plenum Press, New York (1974), p. 47. W. L. Pope, L. H. Smith, and G. F. Srnoot, U. S. Patent 4027494 (1977). D-6 A MAINTAINABLE SUPERCONDUCTING MAGNET SYSTEM FOR TOKAMAK FUSION REACTORS* S. Y. Hsieh, G. Danby, J. R. Powell, P. Bezler, D. Gardner, and C. Laverickt Brookhaven National Laboratory Upton, New York and M. Finkelman, T. Brown, J. Bundy, T. Balderes, I. Zatz, R. Verzera, and R. Herberman Grumman Aerospace Corporation Bethpage, New York INTRODUCTION In the search for inexhaustible energy sources, nuclear fusion energy has received much attention in the past decades. Of the various confinement approaches, the Tokamak is by far the most developed fusion concept. The recent achievements in plasma research at major U. S. fusion laboratories promise demonstration of the scientific feasibility of magnetic fusion within the next few years. In the authors' view, the present route towards the design and construction of superconducting magnets for tokamak reactors, although probably satisfactory for experimental devices, will not be acceptable for commercial fusion power plants. One may accept a certain risk of malfunction for a one of a kind, short lifetime, fusion device, but extremely reliable, maintainable systemswill be required for commercial plants. The DEALSmagnetapproach Cl has been proposedas a way to increase the reliability and ease of maintainance of magnetic fusion reactors. The DEALS [demountable externally anchored low stress] magnet system uses demountable superconducting joints, so that if any portion of the magnet fails, it can be replaced relatively quickly. In addition, the conductor / support structure assembly is arranged so that the conductors transfer the magnetic forces on them to an external reinforcement structure. The load transfer device and demountable joints are designed so that the reinforcement structure operates at relatively high tensile stress/strain levels, while the conductors operate at relatively low compressive strain levels. This concept has been investigated during the past two years by a joint Brookhaven National Laboratory /Grumman Aerospace Corporation study team. A design study * Work performed under the auspices of the Department of Energy. t Consultant to Brookhaven National Laboratory. 207 208 S. Y. Hsieh et al. of a high field ignition test reactor (HFITR) has been carried out, with preliminary experiments on small-scale demountable joints. The conclusion isthat demountable superconducting magnet systems appear feasible for Tokamak fusion reactors. Extensive development work is required, however, before practical large magnet systems can be designed and constructed. The most important need appears to be for further experiments on the mechanical and electrical properties of relatively large scale, prototype, superconducting, demountable joints to establish a data base for design and construction. The balance of this paper describes the latest design approach for the DEALS magnet involving movable pressure-contact, superconducting joints and experiments on such joints. It should be pointed out that these latest design efforts, although they represent improvements over earlier concepts, should not be regarded as the final optimum approach for a demountable magnet. [ 2] CONCEPT OF A MAINTAINABLE, DEMOUNTABLE, SUPERCONDUCfiNG MAGNET SYSTEM The tokamak fusion reactor is a very complex system consisting of a TF magnet, poloidal coils, plasma chamber, etc., all interlocked together. If any component of this system fails, it will be almost impossible to service or to repair the component without making a major disassembly of the highly radioactive reactor. At a minimum, there will be a prolonged plant shutdown of many months, and it may not be feasible at all. In this event the capital investment in the reactor will be lost. In the DEALS magnet concept, TF magnet coils are formed from removable coil segments. These segments can be mass produced at a central facility and then shipped to power plant construction sites for joining. Figure 1 shows a cross-sectional view of a typical DEALS conductor assembly inside a coil case. The conductors are wide (83 cm), thin (0.8 cm) plates of copper MAIN JOOLANT CHANNEl 0.8cm VOIO CERAMIC SIOE INSUL. ClEARANCE GAPS 0.3 cm CERAMIC INSUl. (491 TYP 0.8 cm CONDUCTOR (501 TYP T Fig. 1. Coil case design. Maintainable Superconducting Magnet System for Tokamak Fusion Reactors 209 /'... ENGAGE EITHER THE l r WAY AS SHOWN OR -====7! f- HORIZONTALLY DEPENDI NG ON DESIGN PRETINNED SURFACES Fig. 2. Schematic of idealized demountable joint. Not drawn to scale. with a transposed superconducting braid at the midplane. The conductor is formed by soldering the superconducting braid between two copper plates. Coolant grooves are arranged on the conductor surfaces for heat transfer to a liquid helium bath. The conductors are cryostable with maximum heat ftuxes in the range of 0.3 to 0.4 W /cm2 when all the current ftows into the copper stabilizer. The conductors in each coil segment are typically several meters long. At the ends, one-half of the copper stabilizer is milled away to form a region where the demountable jointwill be made when the coil segments are put together to form the complete coil. The current passes from a given conductor to the next by transfer through the overlapping joint area, which is on the order of 3 to 4 x 103 cm 2 in area. The conductors are arranged so that the completed assembly forms a multiturn coil with the turns in series. Conductors are insulated from each other and from the coil case by ceramic or epoxy-fiberglass plates (see Fig. 1). A schematic diagram illustrating the configuration of the multiple turn joints is shown in Fig. 2 and described in detail elsewhere [ 1•2 ]. The segmented coil approach is a key feature oftheDEALS magnet design and has important benefits in terms of accessibility and maintainability for tokamak reactors. Figure 3 illustrates how all parts of the TF magnet and associated reactor system can be accessed and maintained. Sequence 1 The center tension post is first lowered into place and embedded in the concrete foundation. The tension post can be Iowered into place as one section or each of its 16 components can be separately lowered into place if it is necessary to reduce the maximum Ioad on the overhead crane. The lower collar and the insulator ring are then lowered into a temporary location in the basement and the Iower poloidal field .. \ • \y: . EF/OH COIL RING ASSY ~ : · , lf' • :·~~~~·. ~ING INSULATOR t . .i ._I I I •, • - . - . "l .• ) 2 \IN80 LWR TOROUE SUPPORT STRUCTURE MAGNET LEG " " VEATICAL INBO INBD VERTICAL TOROUE _-I/SUPPORT STRUCTURE ;j i -I Iu__~ II .I · ·.. BUCKING COLUMN/INNER 'OROUE RING ASSY • INSULATO~ LOWER TOAOUE PLATE/ BLOCKS POLOIOAL FIELD COILS ' I, •I ' ~ ... • t. , POST ASSEMBLY SEOUENCE- 1 CENTE~ uPR TO~OUE 'LATE 4 UPR OUTBO TOROUE SUPPO~T STRUCTURE UPR OUTBD MAGNET LEG 5 OUTBD VERT ICAL MAGNET LEG OUTBO VERTICAL TOROUESUPPOATSTRUCTURE VACUUM VESSEL SHIELDING :""' "Q 1:1" r... ~ :< s I I 1 ~ ( -'-------- ( ~· ~ ____,... . ~ . < _j ., .> - \ \ Fig. 3. Total assembly sequence 1 to 6. lWR OUT80 MAGNE T lE G FINGER JOINT AAEA VESSEl iLWR IN PLACIEI 3 P'REASSEMBlED VA.CUUM I OUTBO STRUCTURAL SUPPORT ASSV 6 I-I~ INSUlATOA BLOCKS I UPR INSULATOR RING UPA AlETA lHING AING N .... .... I ~ :::1 f 111" ä =- ~ !! i fi i a 8' ; i ~ ii' J :: 212 S. Y. Hsieb et al. coils and lower torque plate/insulator blocks are seated above the lower collar. This entire assernbly is later repositioned during the third step in the assernbly sequence. Sequence 2 The bucking colurnn and inner torque ring assernbly next is lowered into its final position. The equilibriurn field and ohrnie heating coil ring assernbly is then placed between the bucking colurnn and the tension post. The inboard vertical, and lower torque support structures are assernbled and each of the 16 vertical inboard rnagnet legs is then inserted into its channel and locked into place as shown above. Sequence 3 Each of the 16 lower inboard rnagnet legs is insulated in place and the rnating finger joints of the rnagnet segrnents are connected. The heliurn and vacuurn Dewars are then connected. At this point the lower outboard poloidal field coils are installed followed by the lower outboard torque support structure. Each of the 16 lower outboard rnagnet legs is then positioned and the rnating finger joints connected. The lower collar assernbly is now raised frorn its ternporary position to its final position and locked into place by the lower retaining ring followed by the vacuurn vessel, which can be preassernbled and tested off-line if desired. Sequence 4 At this stage the vacuurn vessel shielding is installed and the outboard vertical rnagnet legs can now be placed in position and the rnating finger joints connected. Sequence 5 The fifth stage in the assernbly procedure involves the installation of the upper outboard and upper inboard torque support structure. Each of the 16 upper outboard rnagnet legs is put in place and the rnating finger joints are connected. This sequence is followed by lowering of the 16 upper inboard rnagnet legs into place. Since these are the last of the rnagnet legs to be installed, rnating finger joint connections are required at both ends of the rnagnet leg. The last step in sequence 5 is then to place the upper torque plate and insulator blocks in the assernbly. Sequence 6 The upper poloidal field coils, upper insulator ring and upper collar are now lowered into position. The upper collar assernbly is then locked into place by the upper retaining ring. This sequence rnirrors the lower collar assernbly in sequence 3. The outboard insulator blocks are installed at this point and each of the 16 outboard structural support assernblies is wheeled into place along tracks. After each of these assernblies is properly aligned, the assernbly sequence is finally cornpleted by inserting the upper and lower pins as shown above. LOW-TEMPERATURE MOVABLE JOINT APPROACH Various types of conductor joints have been used in rnagnet coil applications, including both soldered and pressure contact joints. Both types have been successfully used for roorn ternperature copper and low-ternperature superconducting coils. However, these joints are usually held rigidly together by bolts, reinforcernent, and structure supports. Low-ternperature rnovable joints have been investigated at LASL and MIT [4 ]. e1 Maintainable SuperconductinK Mapet System for Tokamak Fusion Reac:tors 213 A soldered type of demountable, superconducting, movable jointwas proposed in the first DEALS studies C]; while feasible in principle, soldered jointswill require an actively controlled type (e.g., hydraulic pistons) of Ioad transfer device between the coil segments and the external reinforcement structure. A movable pressure contact type joint is now favored as the best approach and forms the basis for the HFITR DEALS design eJ and the assembly sequence shown in Fig. 2. The overlapping regions at the ends of adjacent conductor plates are pressed together with a modest clamping pressure to establish good electrical contact, but are free to move slightly (-1 cm) relative to each other. This eliminates the need for actively controlled Ioad transfer devices between conductors and the external reinforcement structure. lnstead, simple, passive, low thermal conductivity, blocks can be used; the movable joints then permit the conductors to move to accommodate differential thermal and mechanical movements between conductors and the reinforcement structure. The heat leak through the passive support blocks is quite low and permits the use of an external warm reinforcement structure. Such a structure should be substantially eheaper than a cold reinforcement structure. Current transfers from one conductor to the next across the ftat overlapping joint regions. The joint surfaces must provide adequately low electrical resistance and permit small, slow motions without degrading either mechanical or electrical properties. Experiments on properties of movable pressure contact joints are described in the following section. Two design approaches have been developed for applying pressure on movable pressure contact type joints and these are described in the remainder of this section. lt should be remernbered that no tensile Ioads are carried by conductors (other than those resulting from frictional forces associated with joint surfaces, which are small), and that magnetic forces aretransferred to the external reinforcement structure. Relative joint movement is allowed during the following conditions: (1) during cooldown or warmup while clamping pressure is not required; (2) under controlled clamping pressure to achieve tolerable contact resistance when the magnet is energized or discharged in normal operation; and (3) during a quench or emergency shutdown situation that requires quick release of the clamping pressure. The design approach shown in Fig. 4 is a self-activating clamping and/or declamping mechanism which takes advantage of both the Lorentz force which pushes the magnet coil segmentoutward and the support structure element acting as an elastic constraint which confines the magnet coil and pushes it inward. The ramp mechanism (which might be curved for a desirable controlled clamping pressure) is designed to utilize both the outward Lorentz force for clamping the joint and the inward hoop force from the structure for declamping the joint. This will satisfy the requirements mentioned above in (2) and (3). The movable part of the ramp mechanism is designed to develop the initially required joint pressure. This can be adjusted on each individual joint aftercooldown but before the magnet is energized. A second design approach acts in a similar way to the first design except that it is sensor activated. The activator cylinders will operate in the direction (based on the signal) that either exerts or releases pressure on the joint. This is a simple screw thread mechanism. In the magnet start-up sequence, an initial preloading of the joints is applied through the adjustable movable ramp mechanism to ensure sufficient contact area and a low-enough resistance to permit the magnettobe energized. The current is then built up to operationallevel and the 12 R Iosses are monitored to ensure that they are S. Y. Hsieh et al. 214 MOVABLE RAMP MECHANISM (a) n...;;;_----,~ MOVABLE RAMP MECHANISM PRESSURE PLA TE SUPPORT HOUSING PRESSURE----~~---1/ PLATE T/F COIL FINGER JOINT AREA SECTION AA (b) Fig. 4. (a) Conceptual joint design (baseline). (b) Conceptual joint design (baseline). Maintainable Superconducting Magnet System for Tokamak Fusion Reactors 215 within bounds. As the current increases, the electromagnetic forces on the conductors and joints increase and the conductors are permitted to move slightly in the contact region to allow the conductor and joint stresstobe transferred to the external support system. This is accomplished automatically by the ramp design which will exert increasing clamping pressure on the joint when the coil moves radially outward, taking into account the requirement of the tolerable contact resistance and degree of joint movement at every stage of the charging process. During discharge of the coil under normal or emergency conditions, the joint clamping pressure will be released in the first design by the hoop force in the structure when the coils are pushed inward or in the second design by the sensor-activated screw mechanism. PRELIMINARY EXPERIMENTAL RESULTS Since a movable joint is very desirable for demountable superconducting magnets, it is important to demoostrate the feasibility of this joint concept. A small experimental test rig, shown schematically in Fig. 5, was set up to obtain information on some fundamental parameters concerning this type of joint. The various components designed for these test purposes can be identified in the actual equipment shown in Fig. 6. Experimental measurements were made on a set of small sample movable joints (4.74-cm 2 contact area), shown at the bottom of Fig. 6, to determine electrical resistivity as a· function of surface type and contact pressure, friction coefficient as a function of surface type and contact pressure, and effects of surface motion on joint electrical and mechanical properties. Fig. 5. Test rig for simulating movable contacts. l16 S. Y. llsieh et al. Fig. 6. Experimental setup and sample movable joint. A simple experimental setup (Fig. 6) permits testing the joint at various pressures and current densities using small areas of contact surface which can be changed as desired, either by changing the contact size or material type. Surface conditions between the contacting surfaces can also be changed. During a given contact test, electrical resistivity and friction coefficient were measured both statically and dynamically (i.e., when the upper and lower contact surfaces moved relative to each other), as a function of contact pressure. First, contact pressure was monotonically increased from the low end of the range to the upper, with measurements taken at a number of intermediate loading conditions. Measurements were then taken as a function of monotonically decreasing contact pressure; this was followed by another set of measurements under increasing pressure conditions. Several additional Observations can be derived from these experiments. First, the resistance of the joint is essentially the same whether or not the joint surfaces are moving relative to each other. This indicates that the DEALSjoints should be able to move during Operation without adversely affecting performance, particularly since the rates of relative motion for the DEALS application will be very small compared to the rates in the experiments. Second, although joint resistance decreases with increasing contact pressure, it remains relatively low at low contact pressures, i.e., several hundred psi (-2-3 MPa), particularly for the indium coated joints. This indicates that although moderate clamping pressures [e.g., -600 psi (40 MPa)] may be desirable during the steady state, fully energized period to keep joint resistance low, reduced clamping pressures Maintainable Superconducting Magnet System for Tokamak Fusion Reactors 217 [e.g., -300 psi (20 MPa)] during the period of magnet energization or discharge should not result in any significant heating problems. Third, the joint resistance is much greater (one to two orders of magnitude, depending on pressure and surface type) than the resistance of a perfect joint with no interfacial resistance. The reason for this is not certain and needs further study. It may relate to imperfect contact between surfaces or to some thin poorly conducting film on the surfaces. lt is clear that the softer surfaces at the joints, i.e., those with indium coatings, have lower resistances than those with hard surfaces. The thickness of gold on the gold-plated copper surfaces is too small to affect their mechanical conformity. No galling was observed on the thin indium- and gold-coated surfaces, even when high contact pressures [i.e., -1000 psi (70 MPa)] were applied. It appears likely that softer surfaces, or a combination of a soft and hard surface will be more desirable for joints than hard surfaces. Joint resistances were derived from current and voltage measurements on the joint. The voltage measurement was compared with the voltage across a calibrated standard resistance at the same temperature. A calibrated Ioad cell was used to measure contact pressure on the joint. The friction coefficient was then derived from the contact pressure and the applied torque as the joint was moved. The maximum rates of movement between the joint surfaces was estimated tobe -1 cm/s, but no precise measurements were taken. There was no change in joint resistance during movement for the range of 0 to -1 cm/s. For service in aDEALS magnet, rates of joint movement will be very small, on the order of 10- 5 cm/s during normal charge or discharge of the magnet. Four types of contact surfaces were tested: (1) thick indium(-20 mil)-coated copper surfaces; (2) thin indium(- 3 mil)-coated copper surfaces; (3) thin gold-plated copper surfaces; and (4) bare copper surfaces. Figure 7a shows the measured resistances for the two indium surfaces, and Fig. 7b shows the measured joint resistances for the gold-plated copper surfaces. Results for the bare copper surfaces arenot shown; typically, resistances for the bare copper joint were a factor of 2 to 3 higher than those for gold-plated copper joints. The thin indium surfaces give the lowest joint resistance. If the measured value for this surface is used to calculate / 2 R joint heating fortheDEALS HFITR magnet system, the total heating is approximately ten times greater than the value projected in the study. The study value was estimated to be 1 kW at 4 K, which with a refrigeration factor of 326 kWe/kW (4 K), required a refrigerationinputpower of 0.32 MW. Using the experimental electrical resistance values found forthin indiumcoated copper surfaces, total joint heating for the complete TF system would be 10 kW at 4 Kor 3.2 MW for the refrigerationpower input. Adding in the additional refrigeration for the other thermal inputs (current Ieads, passive support blocks, eddy current heating, etc.) the totalrefrigerationpower input is 4.4 MW. This appears to be quite acceptable. Considering that a 1000 MW fusionpower plant will not be much larger than the HFITR, the refrigerationpower input for aDEALS magnet system represents less than 0.1% of the plant output. However, in a better controlled experiment, it is likely that joint resistance can be substantially reduced by further optimization of joint surfaces. The measurements of friction coefficient for indium (Fig. 8a) and gold-plated copper (Fig. 8b) surfaces indicate that the friction coefficient is relatively uniform with contact pressure, and that the barder surface exhibits somewhat lower friction coefficients than the softer one (-0.4 vs. 0.5). The friction coefficient for the indium surface is quite acceptable, however, in terms of aDEALS joint. S. Y. Hsieh et al. 118 • STt.TIC 0 OYNAMIC 3.7 t CXlNTIICT AREA • 474 al0' 4 m1 fXP I IIIIITIAL INCREASIIIIG PRESSURE CYCLE V> ::1: r 03 0 0 u i . w u z 0------ 02 "' >- V> ;;; !I! THICI< INDIUM 13·2.3 ·78) -~ a:: o--------<:l 0 .1 • STATIC V> :::< r 086 0 DYNAMIC f 0 0 a:: u CXlNTIICT AREA • 4 74 al0' 4 m1 :i' w~ 02 ~ ~ THIIII 1NDIUMI4· 20 ·78) :! V> ;;; ~ 01 1000 4000 3000 20CIO APPL IED LOAD, NEWTOIII (a) 176 t RUIIII RUN 2 RUIII3 "':::<r INITIAL INCRE ASING PRESSURE CYCLE DECREASING PRESSURE CYCLE INCREASING PRESSURE CYCLE • STATIC 0 DYNAM IC 05 CXlNTIICT AR(A• 474 a10' 4 m1 0 0 04 a:: u i ~ t:l z :! 03 "' 02 ~ a:: Dl 1000 2000 3000 APPLI[D LOAD, NEWTON 4000 (b) Fig. 7. (a) Contact resistance as a function of applied clamping pressure (indium-plated surfaces). (b) Contact resistance as a function of applied clamping pressure (gold-plated surfaces). Maintainable Snperconducting Magnet System for Tokamak Fusion Reactors 219 I ~ 1 z 8 INDIUM ON INDIUM (4·20·18) 0 f= u er ...0... 0 0 llt CYCLE·UP 2 n4l CYCLE ·DOWN 1- z w Q ...... w 0 u .2 .I 0 1000 2000 4000 3000 APPLIEO LOAD , NEWTON (a) CONTACT AREA = 4 74 x I0" 4 m1 z 0 i= !::! ......0 0: .4 1- z .... u o I st CYCLE • UP .3 0 ii: ....0 (.) 2nd CYCLE- DOWN 0 3 rd CYCLE ·UP 1&. .2 .I 0 1000 2000 3000 4000 APPLIEO LOAD, NEWTON (b) Fig. 8. (a) Coefficient of friction as a function of applied clamping pressure (indium-plated surfaces). (b) Coefficient of friction as a function of applied clamping pressure (gold-plated surfaces). ZlO S. Y. llliell et al. Tbe eflect of joint size on electrical and mechanical properties was not investigated. DEALSjoints will have approximately 103 greater contact area than the small-scale tests described here, and a future development program should examine the eflect of scale. It is likely that if there is any eflect, it will be related to the ability to manufacture large ftat joint surfaces and maintain uniform pressure over this large area. Tbe eflect of magnetic field on joint properties was not investigated either and would have to be examined in any development program. Tbe results of the movable joint tests give strong indication that large-scale movable joints can probably be developed. Tbe next step toward the development of the DEALSmagnet concept is to build a small prototype coil containing several turns of about 1-m length with movable joints. Operation of such a prototype coil would demoostrate the practicality oftheDEALS system on a sufficiently large scale that it could be considered as a viable alternate to the mainline wound coil approach. SUMMARY AND CONCLUSIONS Tbe DEALS magnet concept has significant implications and potentially important benefits for magnetic fusion reactors. DEALS magnets are expected tobe readily maintainable if failures occur, they can be demounted to improve accessibility to other reactor systems (blankets, beam lines, etc.). Tbeir capability to operate at low conductor stress should improve reliability. Magnet components can be made, produced, and assembled at the construction site with minimum field work-elimination of winding stress allows the use of brittle superconductors and insulators. High-current conductors can be employed for rapid energy extraction capability. Pulsed field losses are low, even for heavily stabilized conductors. Design studies indicate that adequate structural support should be readily achieved, practical conductors appear fabricable, cryostable Operation is feasible, and estimated refrigeration requirements are reasonable. Tbe primary technical issue for the DEALS concept appears tobe the electrical and mechanical feasibility of the demountable superconducting joint. Design studies of an attractive joint option, the movable pressure contact type, have been carried out along with experimental measurements of the electrical and mechanical properties of smallscale movable joints. Tbis type of joint allows current tobe carried while overlapping joint surfaces move slightly to accommodate differential thermal and mechanical movements in the conductor/support structure assembly. Mechanisms for applying moderate, adjustable clamping pressures to multiturn joints have been devised. Experiments indicate that the electrical resistivity of such joints is not aflected by relative motion of the surfaces, that joint / 2 R heating is sufficiently low to yield practical refrigeration requiremeqts, that frictional coefficients are reassurably low, and that joint surfaces are not adversely aflected by the applied clamping pressures and relative motion. On the basis of the studies and experiments, it is concluded that the DEALS concept appears feasible. More work should be carried out, including tests of larger joints to develop the engineering base for the design and construction of large-scale systems. ACKNOWLEDGMENTS The authors would üke to thank the foUowing people for their contribution to this paper: J. Jackson, J. Allinger, and J. Weisenbloom for their designing of the test rig and making of the experimental measurements on a series of sample joints, S. Majeski for the drawings, and P. Walton for the typing. Maintainable Superconducting Magnet System for Tokamak Fusion Reactors 221 REFERENCES 1. J. Powell, S. Y. Hsieh, and J. Lehner, "DEALS: A Demountable Externally Anchored Low Stress Superconducting Magnet System for Fusion Reactors," BNL Rept. 50616 (1977). 2. J. Powell, S. Y. Hsieh, P. Bezler, D. Gardner, M. Reich, C. Laverick, M. Finkelman, T. Brown, J. Bundy, R. Herberman, and C. H. von Keszycki, "A Niobium-Tim DEALS Toroidal Magnet System foraHigh Field Ignition Test Reactor," BNL Rept. 50802 (1977). 3. J. Lindsay and D. H. Whitaker, Los Alamos Scientific Laboratory, private communication. 4. Y. Iwasa, Massachusetts Institute of Technology, private communication. D-7 PROTOTYPE LOW-CURRENT SUPERCONDUCDNGQUADRUPOLE MAGNET FOR FERMILAB'S HIGH-INTENSITY LABORATORY* W. Craddock, R. W. Fast, P. Garbindus, and L. Mapalo Fermi National Accelerator Laboratory Batavia, Illinois INTRODUCDON A low-current saddle-type quadrupole magnet using very similar cable and design criteria as its companion dipole Cl has been built and tested in a vertical cryostat. The low current is achieved by taking the 15 film-insulated strands of a cable and connecting them in series. With currents of approximately 250 to 350 A, it becomes economically possible to use a separate power supply for each of the magnets in the beamline. Magnetparametersare given in Table I. COIL FABRICATION Because electrical insulation must be maintained between the individual strands of the cable rather than just between turns, extreme care must be used in the manufacturing of these coils. These coils have, therefore, been built so that they are wound as tightly as possible without damaging the insulation. They must be eventually compressed with moderate preload or thermal pressures in all directions. The quadrupole field was obtained using the ~ rule [2 ]. However, each of the four layers consist of two coils wound from two separate spools of wire. In order to make the transition between layers, adjacent layers have their coils wound around poles which are 90° apart. Thus the coils of layers one and three and the coils of layers two and four are directly above one another. Figure 1 shows layers three and four. The end turns of layer three are just barely visible under the Kevlar banding. This technique minimizes the number of wire splices. It also reduces the magnetic field in the coil ends by separating the end turns and distributing the fJ component of current in opposing directions. Bach of the coils was wound with 60 lb of tension directly on the bore tube around G-10 fiberglass-epoxy poles. As each half-turn was laid down, it was clamped along its length subjected to a pressure of 30 psi. With both half-turns down, the pressure was increased momentarily to 300 psi. This procedure produced tight * Work sponsored by the U. S. Department of Energy. 212 Prototype Low-Current Superconducting Quadrupole Magnet Table I. Magnet Parameters Cable Spiral cable wrapping Cable dimensions Bare strand diameter Conductor Cu/NbTi ratio Effective coil length Total magnet length ColdboreID Inductance Short-sample current Peak field on conductor Cable current density Maximum current achieved Maximum stored energy Design gradient without iron Design gradient with iron Bore tube Outer support pipe Number of layers Inner coil radius Outer coil radius 15 polyester polyamide-imide insulated strands 0.1-mm(0.004-in.)-thick B-staged fiberglass epoxy tape 2.21 x 8.87 mm (0.0863 x 0.349 in.) 1.02 mm (18AWG) 206 filaments of Nb45 wt.% Ti with a ~-in. pitch 2.9 88.4 cm (34.8 in.) 116.5 cm (45.88 in.) 13.3 cm (5.25 in.) 2.8H 343 A at 4.4 K 4.5T 2.10 x 104 A/cm 2 First run: 323 A (48.4 T/m) Second run: 343 A (51.4 T/m) 165 kJ 53.5 T/m (13.6 kG/ in.) 56.3 T/ m (14.3 kG/in.) 8.76-mm(0.345-in.)-thick 2024-T4 aluminum 17.5 mm (0.688 in.) 6061-T6 aluminum 4 7.56 cm (2.975 in.) 11.30 cm (4.45 in.) .040THK GIO INSULATION .,,.,,.. ,. __ .022 x 3fe KEVLAR WEAVE BAND (1800 LB BREAKING STRENGTH) SHEET Fig. 1. Cutaway view of magnet. Third and fourth layers are shown. 223 124 W. Cnddock. R. W. Fast, P. Garbindus, ud L. Mapalo Fig. 2. Coil winding setup. Hydraulic cylinders clamp on the top brackets. end turns. Figure 2 shows the tooling. Upon winding an entire layer, two pairs of split G-1 0 islands are placed in the vacant 30° between the coils, and the winding fixture is removed. The hydraulic compressing fixture, shown in Fig. 3, is then pinned into the holes of the split islands. Each hydraulic cylinder pulls on every other steel ~and which in turn separates the split island and rompresses the wire on the opposite side. The bottom of the conductor slides on a Teflon-sprayed 0.006-in. G-10 sheet which permanently isolates each layer, while a 0.014-in. Teflon-coated fiberglass cloth protects the top of the conductor from abrasion against the bands during compression. The conductor is compressed, neglecting friction, with 2700 psi in the azimuthat direction and 500 psi in the radial direction. At this point the magnet is baked at 380 K for 2 hr. The heating step was initially included to eure the B-staged epoxy tape to insure a morerigid coil package. Thetapethat was used bad very little epoxy, but it was found that the heating procedure compressed each quadrant of conductor by an additional 0.150 in. on the average, probably from a change in the coefficient of friction. During heating, anodized aluminum end saddles are driven against the conductor, preloading the end turns with approximately 500 lb. After the magnet has cooled, shims are tapped into the gap of the split islands. Conductor placementwas within -0.050 in. One at a time, each 6-in. section of compression bands and hydraulics are removed as the magnet is banded with a 1-in. pitch of Kevlar-49t braid. The 0.022 x 0.375-in. braid has a nominal breaking strength of 1750 lb and provides an averageradial pressure of 150 psi. The Kevlar is subsequently relaxed during the preloading of the outer support pipe. Its function is to provide an overall tight conductor package, maintain radial tolerances during winding, and to provide the helium cooling channels during operation. The last fabrication step is to apply the final preload with countersunk stainless steel bolts in the outer support pipe. A radial preload of 700 psi was chosen for the first run and 1000 psi for the second. In addition, four large aluminum reinforcing rings were added outside the aluminum support pipe to reduce the bending moment in the bolts for the second run. t Aramid fiber produced by duPont. Prototype Low-Cnrrent Superconducting Quadrupole Magnet 225 Fig. 3. Compression fixture . Legend: (1, 2) steel bands; (3, 4) 6-in.-long steel blocks; (7) bronze bushing; (8) hardened steel shaft to resist bending moment; and (10) hollow core hydraulic cylinder. STRUCTURA L ANALYSIS Under a biaxialload of 1500 psi, the conductor was measured to have a Young's modulus of 1.5 x 106 psi. When the conductor was subjected to a constant pressure of 2160 psi, the estimated thermal strain from liquid nitrogen to liquid helium temperature was measured as 0.0038. These two values were used in all subsequent analysis. Stresses in the conductor from preloading, thermal contraction, and electromagnetic loading must be considered. The principal preloading mechanism is the bolts in the outer aluminum support pipe. If the conductor has not been wound tightly up to this point, the radial pressure cannot preload the conductor in the azimuthal direction. For calculations, it is assumed that the conductor tends to 116 W. Craddock, R. W. Fast, P. Garbindus, and L. Mapalo behave as a fluid and nearly all the radial pressure from the support pipe is transmitted through to the bore tube. Thermal stresses arise from the difference in thermal contraction between the G-1 0 island/conductor combination and aluminum. The estimated thermal strain of the island/conductor package at 4.2 K is 0.0037. By choosing an aluminum rather than a stainless steel bore tube which the dipole COunterpart uses, radial preload pressure is sacrificed for azimuthat pressure. The conductor can be thought of as being wedged between the islands when cooled to 4.2 K. On cooldown an additional 200 psi radial loading should be gained, although strain gages on the outer pipe indicate a smallloss of preload for the first run. The strain gages were unusable for the second cooldown, but the thermal strain of the conductor is smaller when cooled under greater pressures. The electromagnetic force calculations were provided by Snowdon CJ. When considering the first octant, the radial component of force and the x component of the (J component of force of each conductor are distributed on the outer support pipe. Equations (1) and (2) give the internal bending moments and tensile Ioads in a pipe subjected to uniformly distributed Ioads shown in Fig. 4 as derived from integrating point Ioads [4 ]. M = -wR 2 [(480 /7r + cos 80 - sin 80 ) cos (J- 1]- wR 2 [1 - cos(O- 80 )](8- 80 ) 0 (1) T = -wR[(cos Oo- sin 80 ) cos 8- 1]- wR[l - cos(O- 80 )](8- 80 ) 0 (2) where M is the bending moment, T is the tensile Ioad in lb/in., w is the distributed Ioad in lb/in., and the brackets ( )0 indicate unity for (J ~ 80 and zero for (J < 80 • Fig. 4. Definition of symbols used in (1) and (2). Prototype Low-Current Superconducting Quadrupole Magnet 227 in 0.. -2000 ~ ii: ~ :::1 "'1500 0:: IIJ 1:::l 0 z 0 IIJ !3 1000 ~ IIJ 0:: 0.. Fig. 5. Radialand total electromagnetic loading of outer support pipe at 340 A. 300 0 A superposition of these uniformly distributed loads at various angles approximates the electromagnetic loading on the outer support pipe. The y component of the 8 component is assumed tobe carried by the conductor and bore tube. Figure 5 is the distribution of electromagnetic loads on the outer support pipe. The maximum stress in the aluminum support pipe is 10,000 psi at 323 A. In addition to the redistribution of radial forces, approximately 1700 psi of the tangential preload at the G-10 island is lost on the conductor as it is pinched towards its midplane. The pinch force is 2500 psi at the conductor midplane. Table II is a summary of conductor stresses for the second run. Table II. Estimated Stresses on Conductor (343 A)* u"(r = 3 in.) u"(r = 4.5 in.) u 99 (r = 3 in.) at island boundary at conductor midplane u 99 (r = 4.5 in.) at island boundary at conductor midplane * All units in psi. Preload Cooldown 357 A Total -950 -1000 0 -200 +400 -700 -550 -1900 -1150 -1150 -700 -700 +2300 -20 -2000 -1100 -1100 -600 -600 +2200 -450 -2150 -o -o 221 W. Craddock, R. W. Fut, P. GarWncill8, ud L. ~ ELECfRICAL Figure 6 shows the electrical connections. The dump resistors are center tap grounded with a 100-0 resistor. This reduces the possible valtage to ground in half, but still Iimits the current in case a fault to ground develops during magnet discharge. Furthermore, the magnet was wired in such a manner as to keep sections with the greatest valtage differences as far apart as physically possible. The cable was wired in series with each strand lying next to its two closest electrical neighbors. Each half of the magnet on either side of ground is completely isolated from the other half by islands and a 0.006-in.-thick G-10 sheet between layers. Three out of the 15 strands could not be used. Two pairs of strands developed electrical shorts at the crossover between layers and bad to be wired in parallel. The third wire is a plain copper wire which was used to propagate the quench by shunting current through it. At no time were the two copper wires connected tagether or to ground. When a valtage imbalance is detected, the SCR disconnects the power supply, and simultaneously an auxiliary power supply can energize stainless steel heaters sandwiched between the conductor and the G-10 islands. Three modes of energy extraction were tried and are shown in Fig. 6. At least one of the two copper wires was always left open or connected to the data acquisition system. The safety Ieads were always connected and provide an alternate path for current to bypass the quenching section. COLD~---!--- -3> AMBIENT B I ----,----+1- - - - - - - - V I VOLTAGE TAP TYP V COIL rElTER TYP A I I ------~~----~------.--V COPPER COIL B SCR Q~NCH SWITCH V Fig. 6. Electrical connections. Oose switch A for inductive hot wire technique; close switch B for directly coupled hot wire. The copper coils are actually one strand out of 15 in the cable. Prototype Low-Current Superconducting Quadrupole Magnet 229 TEST RESULTS Figure 7 is tbe quadrupole training curve. Witb tbe firstcooldown 94% of sbort sample was reacbed before it was decided to increase tbe preload and add four reinforcing rings. Before tbe preloading was cbanged, all tbe bolts in tbe outer support pipe were totally relaxed. From previous experience witb low-current dipole racetrack coils and from tbe Energy Doubler program [5 ] , it is expected tbat little if any training from tbe first run can be retained after a preload relaxation. On tbe second cooldown, tbe sbort-sample Iimit was reacbed witbin tbree quencbes. Thus tbe 1000 psiradial preload gives good training bebavior witb minimum conductor stresses. The original design gradient was not quite reacbed. Had tbe two sborted pairs been available, it is probable tbat 56 T /m witbout iron could bave been reacbed. It is especially desi'rable witb tbis style of low-current magnet to minimize tbe voltages during a quencb by selecting tbe smallest value energy dumping resistor wbicb is consistent witb avoiding a wire burnout. Stainless steel strip beaters along tbe islands, and an inductive and direct coupled bot wire tecbniques were all tested to measure tbeir effects on quencb propagation. If botb copper circuits are open, ac Iosses in tbe wire are ignored, and quencbing does not occur; tbe square root of tbe joule beat dissipated in tbe external resistors must be proportional to tbe magnet current prior to firing tbe SCR. This is curve A in Fig. 8. Curve B, called coasting witbout beaters, refers to tripping tbe SCR prior to a spontaneous quencb. Above tbe 100-A Ievel some of tbe energy is released in tbe 4.2-K environment. Near tbe sbort sample Iimit, 35% of tbe energy is dissipated internally. The effectiveness of tbe beaters is tbe difference between curves B and C. The stainless steel beaters bave a typical surface beat ftux of 18 W/cm 2 witb an average beat input of 3500 J and a maximum temperature of 120 K as measured witb cbromel-constantan tbermocouples. Curve Eis tbe inductive bot wire tecbnique. This circuit was found tobe not 400 - 350 U) 0.. ::!' <{ ...z w 250 a: a: CRYOSTAT PRESSURE • -3 PSIG x -0 PS IG ::> u ...w 200 z C> <{ ::!' 150 100 50 Fig. 7. Quadrupole training curve. Curve A , 700 psi preload ; Curve B, 1000 psi preload. 5 10 15 OUENCH NUMBER 20 130 W. Craddock, R. W. Fast, P. Garbincius, and L. Mapalo -~ 13 ~ :> ~ 12 "' · II -~ "'a: ~ i3"' a: .J "'z ffi I- 10 9 a 7 X "' 6 ~ ß 0. 5 B 4 :I; ,... ~ 3 "'z "' 2 0 100 200 300 400 MAGNET CURRENT . AMPS Fig. 8. Energy extraction at room temperature vs. magnet current. (A) Theoretical 100% efficiency; (B) coast witbout beaters; (C) coast witb beaters; (D) quencb witb beaters; (E) inductive bot wire tecbnique witb beaters; (F) directly coupled bot wire tecbnique witb beaters. Letter "0" denotes a spontaneaus quencb. nearly as etfective as the directly coupled copper wire, curve F, where roughly half the total energy is absorbed by the coil at 250 A. Changing the polarity of this copper loop made little ditference in the quench propagation at all current Ievels. From 40 to 60% of the energy which is removed at room temperature is dissipated in the 1.43-0 current limiting resistor of the copper loop which is to be expected from the ratio of the resistances. Obviously, if both copper loops were used, quench propagation would have been further improved. Basedon adiabatic conditions, the integral JI 2 dt can be equated to a maximum bot spot temperature in the superconducting windings [6 ' 7 ]. A measured value of RRR = 70 was used in these calculations. Under normal conditions a 337-A coast without heaters bad the highest theoretical bot spot temperature of 250 K. A 200-K hot spot was the worst case for any of the current driven hot wire runs. The Iow external resistance of 0.83 0 of this circuit caused this comparatively high temperature but reduced the peak voltage to ground by 60%. At 273 A with both copper loops open, the SCR and stainless steel heaters failed to trip during a quench. The magnetwas self-protecting and experienced a 275-K bot spot in the quenching loop. The safety Ieads probably saved the magnet from a burnout. The thermal stability of the magnet was investigated by powering the stainless steel heaters and one of the copper loops. For one of the stainless steel heater tests, the magnet current was set at 308 A. With a surface heat ftux of 0.03 WI cm of cable, 0.057 J I cm induced a quench. Next the current was adjusted to 333 A, 97% of short sample. This corresponds to a thermal temperature reserve of only 0.15 K [8 ]. Fifty-three watts of heat were dissipated in one-half the magnet by one of the copper wires for 6i min before quenching occurred. In the highest field region of the magnet, the power density was 0.0027 Wlcm. The relatively porous nature of this conductor allows this stable performance even with currents very close to the short-sample Iimit. Prototype Low-Current Superconduding Quadrupole Magnet 231 FUTURE Twenty-six full-length, 10-ft quadrupoles are eventually required. To reduce the costs, the full-size quads will be two-layer, 4i-in. clear bore magnets using a copper to superconductor ratio wire of 1.8 : 1 or smaller. There is no plan to place this prototype in the High Intensity Lab. Beam quenching sensitivity will be tested with one of the low current dipoles. ACKNOWLEDGMENTS The authors wish to thank J. Satti, P. Mazur, B. Cox, S. Snowdon, J. Heim, and P. Mantsch for their suggestions. A special thanks goes to S. Anderson who was in charge of magnet construction and who helped perfect the many special techniques. They also wish to acknowledge R. DeNeen, J. Guerra, D. Garcia, S. Dochwat, S. Tonkin, and L. Sawicki, who helped with the construction and test setup. HEFERENCES 1. B. Cox, T. Dilman, P. H. Garbincius, L. Kula, P. 0. Mazur, J. A. Satti, A. Skrabaly, and E. Tilles, IEEE Trans. Magn. Mag-15(1):126 (1979). 2. A. Asner, Proc. 1968 Summer Study on Superconducting Devices and Accelerators, Part III Brookhaven National Labaratory Rept. 50155 (C-55), Upton, New York (1968), p. 866. 3. S. C. Snowdon, private communication # 102577-0930. 4. R. Roark and W. Young, Formulas for Stress and Strain, McGraw-Hill Book Company, New York (1975), Table 17. 5. W. B. Fowler, P. V. Livdahl, A. V. Tollestrup, B. P. Strauss, R. E. Peters, M. Kuchnir, R. H. Flora, P. Limon, C. Rode, H. Hinterberger, G. Biallas, K. Koepke, W. Hanson, and R. Brocker, IEEE Trans. Magn. Mag-13(1):280 (1977). 6. P. Eberhard, M. Alston-Garnjost, M. A. Green, P. Lecomte, R. G. Smits, J. D. Taylor, and V. Vuillemin, "A Burnout Safety Condition for Superconducting Magnetsand Some of Its Applications," Lawrence Berkeley Labaratory Rept. LBL-7272 (September 1977). 7. B. J. Maddock and G. B. James, Proc. IEE 115(4):543 (1968). 8. J. Allinger, G. Danby, H. Foelsche, J. Jackson, D. Lowenstein, A. Prodell, and W. Weng, IEEE Trans. Magn. Mag-15(1):119 (1979). · D-8 SUPERCONDUCfiNG MAGNETS OF TUE BIOMEDICAL FACILITY AT SIN J. ZeUweger, G. Vecsey, and I. Hoi'Vath SIN Swiss Institute for Nuclear Research Villigen, Switzerland INTRODUCfiON Negative 1r-meson therapy has gained more and more interest in the past few years because of the unique absorption characteristics and depth dose behavior of these particles with a high linear energy transfer (LET) near the desired region. The technique could be superior to conventional X-rays, and may constitute major progress in tumor therapy. Furthermore, three-dimensional tumortreatmentwill be possible. A test facility for tumor irradiation therapy with negative 1T mesons is under construction at SIN [ 1]. Superconducting ac magnets for tumor scanning [2 -4] are necessary and represent a major part of such a facility. Similar to the Stanford double torus spectrometer CJ, the facility uses 60 identical beams, with 60 pancake coils and 60 slits. A design ftux of up to 2 x 109 1r/s will be concentrated on the tumor tissue. More design details are given elsewhere [4 '6--8]. Two prototype coils were tested. The first torus with its 60 pancake coils is presently under test. Results are presented in this report. THE PION APPLICATOR The applicator (Fig. 1) essentially consists of two toroidal magnets on a common axis defined by the primary proton beam. The toroidal field is generated by superconducting pancakes (60 per torus), subdividing the pion ftux into 60 individual beams of common central momentum. The momentum width and intensity of each beam is controlled individually by a variable slit in the dispersion plane. The magnets are suspended in a common vacuum chamber; optical alignment can be established in the cold state. Forwardneutronsand protons are shielded from the patient by a large cylindrical steel block (140 Mp) inside the chamber. Thermalradiation shields at 90 K cover the inner walls of the tank and a proton collimator should protect the entrance coils from radiation darnage and excessive nuclear radiation heating. The pion beams leave the vacuum through 60 stainless steel windows and enter the patient after crossing a monitor chamber assembly. The position of the patient, the pion momentum (i.e., magnet current) and the slit position are controßed by a microprocessor according to a previously calculated treatment plan. Z3Z Superconduding Magnets of tbe Biomedical Facillty at SIN slit system vacuum chilmber p;~t ient 233 tank Fig. 1. Cross-sectional view of the pion applicator. SUPERCONDUCI1NG MAGNETS Patient size, treatment volume, maximum treatment time, available beam intensity, pion physics, and the beam optics formed the basis of the magnet design. The overall size of the vacuum chamber including the magnets is based on the following requirements: 1. The patient tank bad to be large enough to allow adequate scanning movements of the patient in three dimensions and to incorporate additional equipment. 2. The proton and neutron shield bad to be thick enough to absorb the spent 600-MeV proton beam. 3. With the design primary proton current of 20 p,A, an acceptance angle of 1 steradian at 60° pion production was necessary to stay within the desired treatment time for a treatment volume of 1liter. 4. An average pion lifetime of 26 ns at rest limited the overall path length. 5. A dispersion-free optical system with a magnification of approximately 1: 1 was desired. First-order calculations fixed the main dimensions of the coils and the important beam parameters such as "curvature x radius." Second-order calculations introduced concave curvatures and slight rotations of the entry and exit regions of the coil relative to the centrat beam trajectory (see Figs. 1 and 2). Results of the calculations are summarized elsewhere [6 ' 7 ]. For convenience in treatmentplan development, the irradiationangle at the target image (tumor) in the patienttank was chosentobe 90°. Patient size, the beam parameter "curvature x radius," and the penetration depth of pions defined the required B field. The desired treatmenttime necessary to scan the J. Zellweger, G. Vecsey, and I. Horvath 234 \ TORUS 2 \ TORU 5 1 Fig. 2. Pancake coils of torus 1 and torus 2. treatment valume defined the scan frequency and amplitude (:fo Hz and 10% af Bmu). The superconducting cable was fabricated af Nb-Ti with a copper stabilizer. The design current selected was 2500 A and was a campramise between small ac lasses, small current transfer lasses, law quench discharge valtage, and convenient size for winding pu~ases. Elaborate conductor design resulted in a high current density af 220 A/mm with a low average ac loss of 1.5 mW/m (or 15 W/torus) at 5.5 K and 2.8 T. Summarized data are given in Table I. The coil frame was constructed of glass-reinforced epoxy material mainly to minimize ac lasses. The frame was designed to withstand the magnetic pressure on the winding (120 kp/cm). In the energized state, a resulting force of 2600 kp (3200 kp for torus 2) acts on each coil. Rivets were needed for reinforcement of the two-piece frame and to ensure an acceptable displacement of the patted winding during operation. The pancake coils in each torus are electrically connected in series. The valtage of each coil is continuously monitared by two detectors (for safety purposes) and Table I. Cable Data Electrical data l c. A/mm2 220 (at 5.5 K and 2.8 T) Im.., A2500 P 80, W /torus 15 (at fo Hz and 0.1 Imaxl Dimensions 4.15 mrn x 2.9 rnrn Cable Copper area, A eu. rnm2 NbTi area, ANbTi> rnm 2 Aeu : ANbTi Filamentdiameter, ~rn Nurober of filarnents 7.5 1.05 7:1 ::530 1400 Superconducting Magnets of the Biomedical Facility at SIN _ 235 ·--1-"'F-rivet I·- FU11.ltJ y- p I copper cooling ~ tu~s IA'III!III-+-+-T-IIIIi• · fiilll=t=t=t:::1\Mf-11'!11r-t--t--t--lllliJ - Nb-Ti- copper w inding · - -1- ili 1-+-+-+-!IU:,l"_, m-+-+-+-m'· I IJIIII--t---r-t-__,11111 I , - COOling ITleSh IIIU-+-+-+-m·tll I, Fig. 3. Cross-sectional view of the superconductor winding. compared to the voltage of the neighboring coil. The difference in the two voltages is sensed and used as a signal for quench detection. Bachtorus can be deenergized in 2 s following a quench and the energy dissipated in a parallel shunt resistor. The discharge voltage rises to 1.5 kV, when the torus is fully energized. The coils and their support rings are indirectly cooled-two separate copper cooling tubes are used (see Fig. 3)-by forced ftow of supercritical helium at an inlet pressure of 10 atm. Efficient heat removal in the winding is maintained by aseparate cooling mesh. The coils are cooled in series. An additional heat exchanger in the helium bath guarantees that a temperature of 4.5 K is maintained in the cooling circuit after each group of five coils. The helium is circulated by a multipurpose helium refrigerator, which considers each torus to be a separate Ioad. Details are given in a separate report [8 ]. Specific data for the coils and the magnets are summarized in Table II. COIL TEST Two pancake coils of torus 1 were chosen for this test. The short sample current for this cable was determined tobe 2800 A at 5.5 K and 2.8 T, 12% above the design current. A 7 -Mp iron double mirror was used to simulate the toroidal magnetic field (see Fig. 4 ). The supportring configuration was simulated using appropriate supports and one bearing. J. Zellweger, G. Vecsey, and I. Horvath 236 Table II. Magnet Data* Coil data Bmax at the winding, T Current, A Turns Ampere turns, A-turns Enclosed flux per turn, V-s Average length, m Magnetdata Magnetic energy, MJ lnductance, H B field x radius, T-m Totalturns Totallength, m Torus 1 Torus 2 2.84 2408 60 144,480 0.707 2.8 2.7 2412 48 115,820 0.99 3.5 3.0 1.0 1.73 3600 10,800 3.5 1.2 1.4 2880 10,800 *Operating temperature, 4.5- 5.5K; pion momentum, 220MeV/c, where c is the velocity of light. - ~ I i iron double I mirror I I ! I \ II I coi ls -- . I r- \ I! . I .i 'l '------------ I l I I I I support > cool ing tube s bearing _j Fig. 4. Test assembly for two pancake coils. Superconducting Magnets of the Biomedical Facility at SIN 237 Prior to the test, the coils, the iron mirror, the supports, and the bearing were cooled down to the circulating helium temperature of 4.5 K. The test results are as follows: 1. The cooldown time of the test assembly, having approximately the same weight as the torus, was about 3 days. 2. Five quenches were observed before the design current of 2500 A was reached: one quench for coill; four quenches for coil2. After an additional quench, 2650 A was attained. The "worst-case" cable temperature was 5 K during an energizing cycle. Training steps are shown in Fig. 5. For coil 2 the current increase per step was constant and showed no statistical distribution, a behavior which has not been explained. 3. Owing to the mechanicallimitations of the coil, the critical current of the cable at 5.5 K could not be determined. 4. Although the ac loss of 4 W per coil was too high, the coil could be swept and ramped at the specified frequency and amplitude. An electrical short circuit in the cooling structure owing to inadequate glass insulation was found to cause this· excessive ac loss. 5. Coil performance remained unchanged even after numerous thermal cycles, energizing cycles, or quenches. The coils withstood the magnetic and thermal stresses without any visible darnage or loss of performance. The performance of the two prototype coils of torus 1 met the design requirements except for the excess ac loss, which was eliminated by a thicker insulation :;{ C> C> "' N -;: C> C> 90 80 70+---;----r---+----~--+----r---+-+-5 4 2 COIL 1 COIL 2 number ol trAining steps Fig. 5. Training steps for the two pancake coils of torus L J. Zellweaer, G. Veaey, anc1 1. Homath layer. With this minor design change, important for future coil production, and with the encouragement of these results it was decided to advance immediately to the full torus test. TORUS TEST Torus 1 with its 60 pancake coils is presendy under test. The short sample currents of the cables used range from 2650 to 3400 A at 5.5 K and 2.8 T. Preliminary results are as follows: 1. The cooldown time of torus 1 was 4.5 days. 2. 2410 A was attained after 56 quenches. The first quench occurred at 1530 A and the 19th at 2000 A. During each energizing cycle, the "worst-case" cable temperature was 4.9 K. After a quench, the average temperature of the quenched coil increased to 20 K. No otber coil except the quench-initiating coil was observed to quench, although the discharge time was 2 s. Twenty-seven coils did not quench during the tests. The plot of the training steps shows an exponential characteristic with a slight curvature. The current changes per training step show statistical behavior as expected. There is no correlation between the coil or the amount of quenches per coil and the short sample current; this supports the theory that training is induced by the mecbanical energy release associated with cracks or conductor displacements since all coils were exposed to identical boundary conditions and were operated weil below the critical short sample current. 3. Alt~rnating current loss of the torus at the specified amplitude and frequency (Bmax = 0.06 T/s) is 10 W, wbich is below the calculated value. The cable maintained its superconducting state even after a fast discbarge with Bmax = 1.4 T / s following a quench, verifying its excellent ac performance. The results presently obtained with torus 1 already fulfill the requirements for patient treatment. However, final results will be reported when available. REFERENCES 1. 2. 3. 4. 5. 6. G. Vecsey, SIN Rept. TM-62-01 (1976). H. Blattmann, Rad. Environmental Biophysics 16:205 (1979). J. Zellweger, SIN Rept. TM-62-02 (1978). J. Zellweger, SIN Rept. TM-65-01 (1978). D. Boyd, H. A. Schwettman, and J. Simpson, NucL Instrum. Meth. 111:315 (1973). R. Frosch and J. McCulloch, SIN Rept. TM-37-003/004/005 (1973). 1. J. Crawford, SIN Rept. TM-37-008-009 (1977). 8. J. Zellweger, in Proc. 6th Intern. Magnet Technology Conference, ALFA, Bratislava, Czechoslovakia (1977), p. 361. E-1 A NOVEL THERMOMETER SENSOR FOR TUE mK REGION USING THE PROXIMITY EFFECT H. Nagano, Y. Oda, and G. Fujü Tokyo University, Tokyo, Japan INTRODUCTION A normal metal which is in good electrical contact with a superconductor shows superconducting properlies because of the proximity effect C· 2 ]. This effect in Cu-clad Nb and Cu-clad Nb-Ti wires has been measured and is thought to be applicable as a new type of thermometer sensor below 1 K [ 3 ' 4 ]. When the temperature of such a wire is decreased, the Nb part of the wire transforms to the superconducting state at Tc, the critical temperature of Nb. If the temperature is decreased further, the Cooper pairs in the Nb leak into the Cu. The leakage distance of these Cooper pairs depends on the temperature and the electron mean free path IN in the Cu. The leakage distance increases as the temperature decreases and as the electron mean free path is increased. Thus, in a weak magnetic field, a Meissner effect is seen in the Cu, caused by the leakage of the Cooper pairs into the Cu. The region of this Meissner effect increases with a decrease in temperature. Resistance of the Cu perpendicular to the Nb wire is expected to decrease as the temperature is lowered because of an increase in the leakage distance. Since the leakage length of the Cooper pairs depends on some function of the temperature, it is possible to use the effect as a thermometer below 1 K. MAGNETIC SUSCEPTIBILITY The magnetic susceptibility of Cu-clad Nb wire has been measured by an ac mutual inductance bridge. In practice, the mutual inductance is proportional to the susceptibility. Thus, the size of the Meissner region can be calculated from the mutual inductance of the Cu-clad Nb wire. Forthis purpose the specimen is attached to the mixing chamber of a dilution refrigerator and the temperature is measured by both germanium and carbon thermometers. At the critical temperature of Nb the change in the mutual inductance is expressed as !l.M0 (Tc), while the !l.M(T) represents the change of the mutual inductance after the transition. The latter change is caused by an increase of the Meissner region in the Cu. The size of the Meissner region is expressed as !l.M(T) {[ (1) p = r 1 + !l.Mo(Tc) - 1 ]t/2 } In this relation p is the distance of the Meissner region extending into the Cu from the boundary of the Nb/Cu interface and r represents the radius of the Nb core wire. The 2.39 H. N...-. Y. Oda, ud G. Fejli .· 1.0 . 30 E ·' "- ... 20 .. 10 .. .. ~0 .~ 00 / 2 3 _LI. ·t5 6 8 T2. K Fig. 1. Temperature dependence of the Meissner region, p, for Cu-clad Nb wire (0.35 mm Cu OD and 0.25 mm Nb 10). Residual resistance ratio of material in upper curve is 90 while for the lower curve it is 55. cylinder of radius (r + p) is the Meissner region. From the experimental evaluation of äM0 (Tc) and äM(T), p can be calculated by use of (1) [4 ]. The temperature dependence of p is shown in Fig. 1. The horizontal axis is plotted as a function of r-t/2. This figure indicates that p is a linear function of T- 112 below about 1 K. The Meissner region, p, becomes larger when the resistance ratio of the Cu is larger. Above 1 K, pisalinear funtion of T- 1 instead of T- 112 • In the region where the electron mean free path IN is shorter than the coherence length ~N in the Cu, p has been calculated theoretically by Deutscher and de Gennes [ 1]. From their calculation, p is expressedas p = ~N [log ~Nf A (0) - 0.116] (2) where A(O) represents the penetration depth at the Nb/Cu interface. The function inside the parenthesis is only minimally affected by the temperature. Thus, the temperature dependence of p is determined mainly by ~N· In the approximate Iimit the relation between ~N and the temperature is expressedas ~N = (hv,.IN/6'7Tk8 T) 112 • This signifies that p is a linear function of T- 112 • In this relation, v,. is the Fermi velocity of Cu. When IN is larger than ~N. the relation between ~N and the temperature is expressed as ~N = hv,./2'7Tk8 T; thus, above 1 K, IN is larger in magnitude than ~N and p shows a T- 1 dependence. The thickness of the Cu cladding is expressedas d, and x represents the distance from the Nb/Cu interface. In the region where x < p, Cu shows perfect diamagnetism. In the region where d > x > p, Cu is in the normal state, and the magnetic ßux can be assumed to penetrate freely. In this case, the magnetic susceptibility of Cu in the region of x < p is -!",, and the susceptibility of Cu in the region d > x > p is essentially zero. Thus, the average magnetic susceptibility of Cu, expressed as Xav• can be given as 1 Vp Xav = - 4'7T Veu (3) A Novel Thermometer Sensor for the mK Region Using the Pro:dmity Eßed 241 20 .-------.-------.-------~----· 0 15 0 0 0 5 0 3 Magnetic Field • Oe Fig. 2. Magnetic field dependence of p at 99.5 mK. The average susceptibility of a specimen whose resistance ratio is 90 can be determined from Fig. 1 as Xav = 10- 2 T- 112 • This value should be compared with the magnetic susceptibility of CMN, which is generally used as a thermometer below 1 K, of 8.5 X 10-4 T- 1• Thus, at 1 K the Xav is about twelve times larger than the susceptibility of CMN. Since the temperature dependence of Xav and XCMN is markedly different, XcMN > Xav at very low temperatures. However, from 1 K to about 10 mK, Xav is larger than XCMN· If Cu-clad Nb wire is used as a thermometer sensor below 1 K, its sensitivity is stilllarger than that of the CMN above 10 mK. Moreover, the thermal contact with the metal is quite good. Because of these characteristics the Cu-clad Nb wire appears as a good temperature sensor for the mK region. MAGNETIC FJELD DEPENDENCE The magnetic field dependence of p is shown in Fig. 2. This plot shows that p is very sensitive to magnetic fields. lt decreases significantly even in a magnetic field of less than 1 Oe. At lower temperatures, p is very sensitive to the field strength andin the presence of a field the temperature dependence of p is not a linear function of T - 172. If shielding from a magnetic field is less than 10 mOe, the effect from the field can be neglected. Usually, the magnetic field in a laboratory, which is mainly the earth field, is about 500 mOe; so, if the IL-metal shield is doubled or tripled, it is not too difficult to reduce the residual field to less than 10 mOe. Such types of shielding are quite common in experiments involving Squids. ELECfK.ICAL CONDUCTIVITY The preliminary measurement of the electrical resistance of Cu-clad Nb is shown in Fig. 3. The preparation of the specimen was as follows. First, Cu-clad Nb-Ti multifilament wire of reetangular cross section was rolled into a thin tape. Then it was cut alternately from both sides at right angles to the direction of the filaments (see Fig. 4). The resistance was measured using the four-probe method. The resistance of such a specimen changed suddenly at the Tc of Nb-Ti similar to the mutual inductance change observed in Fig. 1 at Tc. As the temperature was decreased below Tc, the resistance also gradually decreased. It is thought that this part of the decrease was caused by the proximity effect. The decrease in the resistance does not show the H. Nagano, Y. Oda, and G. Fujii 242 2.0 c: E ." 1.0 .~ ~· ·:. ~· .···t 0 ~0------~ 5--------.~0------~.5 T, K Fig. 3. Typical behavior of the electrical resistance, R, of Cu surrounding many Nb-Ti fine wires. Measuring current is 10 mA dc. T- 112 dependence. There are two possible reasons for this nondependence. First, the distance between each Nb-Ti fine wire varies throughout the sample. When the temperature is slightly below T" some of the Cu is in the normal state surrounding each Nb-Ti wire. Because of the proximity effect, the superconducting region progressively increases in the Cu and the change in the resistance becomes large. With further decrease in temperature, the superconducting region in the Cu continues to increase and finally overlaps with other similar regions. The rate of resistance change with respect to temperature decreases. When the Nb-Ti fine wires arenot uniformly distributed in the Cu specimen, the temperature dependence of the resistance is somewhat difficult to evaluate. The other reason that it does not show the T- 112 dependence isthat the magnitude of the dc current affects the measurement. In the case of Fig. 3 a dc current of 10 mA was used. In this case the superconducting region in the Cu may have been decreased by the magnetic field of the dc current. Usually, with the conventional dc method it becomes difficult to attain sufficient sensitivity. The temperature dependence of the resistance for the same type of specimen measured by an ac bridge using the Squid as a null detector is shown in Fig. 5. In this case a current of 10- b A was used. The temrerature dependence is different from that of Fig. 3, but there is still neither a T- 1 / or a T - 1 dependence; this is probably because of the random distribution of the Nb-Ti fine wires. Fig. 4. Schematic diagram of specimen prepared for the resistance measurement. A Novel Thermometer Sensor for tbe mK Region Using tbe Proximity Etlec:t 243 90 r-------~--------.---------.-------~-----. 0 c: '='~. 0 80 UJ u 0 0 z ,_ <{ 0 "'~ 70 0 a: 0 60 cP 0 0 TE MPERATU RE , I< Fig. 5. Temperature dependence of the resistance due to the proximity effect. Measuring current is 10- 6 A. CONCLUSION The magnetic susceptibility change of Cu-clad Nb (or Nb-Ti) wire, caused by the proximity effect, shows a T- 11 2 dependence as an approximate limit, and is thought tobe applicable as a new thermometer sensor. The temperature dependence of the electrical resistance for such a specimen has been studied. In order to ascertain an accurate temperature dependence it is necessary to utilize a specimen in which the superconductor wires are uniformly distributed. However, even a nonuniform randomly Cu-clad superconductor could be used as a secondary temperature sensor if the resistance and the temperature dependence is known. REFERENCES 1. G. Deutscherand P. G. de Gennes, Superconductivity, Vol. 2, R. D. Parks, ed., Marcel Dekker, lnc. New York (1969), p. 1005. 2. Orsay Group, Phys. Condensed Matter 6:307 (1967). 3. Y. Oda and H. Nagano, J. Phys. Soc. (Japan) 44:2007 (1978). 4. Y. Oda, G. Fujii, and H. Nagano, Jpn. J. Appl. Phys. 18:1411 {1979). DISCUSSION Question by S. W. Van Sciver, University of Wisconsin: Have you accounted for the Kondo effect in the copper? This should have a temperature dependence that is similar to the susceptibility, i.e., xcc 1/TI/ 2. Answer by author: The impurities in Cu of our specimen, Cu-clad Nb, are thought tobe about 10 ppm on a weight ratio. The magnetic susceptibility attributed to the Kondo effect is paramagnetic and its magnitude is about 10- 7 emu at 1 K. In the case of the proximity effect, the magnetic susceptibility is diamagnetic and its magnitude is 10- 2 I T 112 • This magnitude is fairly !arge in comparison with that expected from the Kondo effect. E-2 A SUPERCONDUCTING RF NOTCH FILTER* C. S. Pang, C. M. Falco, R. T. Kampwirth, and I. K. Schuller Argonne National Laboratory Argonne, Illinois and J. J. Hudak and T. A. Anastasio Department of Defense Fort Meade, Maryland INTRODUCfiON Over the past several years major improvements have been made in the quality of high Tc A-15 superconductors. In particular, high-quality thin films of Nb 3 Sn and Nb 3 Ge have been fabricated using chemical vapordeposition C], coevaporation, [2 ] and magnetrOD sputtering techniques. Aside from the potential applications of these materials to ac power transmission, high-field magnets, and particle accelerators, their high transition temperatures might be exploited for use in devices cooled by closed-cycle refrigeration. Most superconducting devices such as the various forms of Josephson junction devices, super-Schottky diodes, cavity and helical resonators, and thin-film bolometers generally require operation at liquid helium temperatures. Low operating temperatures are often necessary to reduce device noise, provide lower resistive losses, or obtain the specific superconducting properties offered by low Tc materials. The inherent disadvantage of low operating temperature is that refrigeration is typically provided by liquid helium baths which must be replenished frequently. In addition, the transfer of liquid heliuni from storage dewars is awkward and unacceptable to a nontechnical user. Small Joule-Thomson expansion liquefier stages added to closed-cycle refrigerators can be used in some applications, but are not readily available. Applications that would allow use of high-quality, high Tc thin-film materials and allow operating temperatures between 9 and 15 K could take advantage of the reliable refrigeration systems which are now commercially available. A preliminary investigation is presented of a radio frequency superconducting notch filter employing thin-film technolo~ which could be used for interference reduction in a communication system [ ]. eJ * Work supported by DOD under MIPR-H98230. 244 A Superconducting RF Notch Filter 245 Fig. 1. Model circuit of notch filter. CIRCUIT DESIGN The circuit used in this investigation is shown in Fig. 1. The filter consists of an inductively coupled superconducting tank circuit which is placed in parallel with the Ioad of the receiver, RL. The mutual inductance between L 1 and L 2 is controlled by the coupling constant, k. The circuit is tuned by varying the capacitor C 2 and the resistive Iosses of the tank circuit are modeled by the series resistance R 2 • For a narrow band of frequencies the filter impedance becomes very small with respect to the Ioad, thereby reducing the strength of the interfering signal at the receiver. The maximum value of the notch depth, D, the 3-dB bandwidth of the notch, tl.f, and the center frequency, / 0 , at which maximum attenuation occurs, are given by [5 ] D(dB) = -20 log (1 + B) 1 Re [ 2 1 ] 112 tl.f = 27T L 1(1- k 2 ) 1 + B- B 2 fo = __!_ [L2C2(1 27T - e)]- 112 (1) (2) (3) where B = ReL 2 /R 2 Lt and R = R.Rd(R. + RL). Equation (1) shows that reducing R 2 leads to a deeper notch if the other circuit parameters are kept constant. lt is also clear that the circuit has the property that the bandwidth and notch depth are independent of the center frequency. In the intended application this is a desirable feature which allows for a very large tuning range. Knowledge of L~. L 2 , D, and tl.f allows one to solve for the value of R2. Thus, the circuit can also be used to study the nature of the superconducting losses. CIRCUIT REALIZA TION The filter circuit was implemented using the planar geometry shown in Fig. 2. The coupling constant, k, between L 1 and L 2 was controlled by the separation distance between plates A and B, and the capacitance, C2 , was controlled by the separation between plates B and C. The present design allows the superconducting tank circuit to be built without the need for any plate to plate connections, thus eliminating contact resistance as a potential source of problems. The planar geometry was chosen so that spottered films of high Tc Nb 3 Sn or Nb 3 Ge could be used. However, for these initial experiments niobium films were used for convenience. Films, about 0.5 #Lm in thickness, were deposited using electron beam and spottering techniques on S-em-diameter single-crystal-sapphire (Ah0 3 ) and fused-quartz (Si0 2 ) substrates. In order to obtain better adhesion between the niobium film and substrate, the latter was held at a temperature of about 300 to 400°C during the deposition process. Scanning electron microscopy showed the film surface tobe smooth on the scale of 0.1 #Lm. 246 C. S. Pang et al. L, A L2 B c c~' c~' FILTER CONFIGURATION Fig. 2. Geometry of circuit elements and their relative position. (Notice the plate-to-plate connections are not required in this design.) The circuit elements were etcbed into tbe films using pbotolitbograpbic tecbniques. The circuit patterns for tbe masks were computer generated and plotted on a . Gerber plotter; emulsion masks were made by pbotoreduction with a x 10 reduction camera. Shipley AZ-1350J photoresist was spun onto tbe films and exposed by a mercury arc lamp using direct contact printing. The etchant used for tbe final development of tbe coil and plate patterns consisted of one part H 2S04 , two parts HF, one part HN0 3 , and four parts H20 by volume. Tbis etcbant was found to work weil on Nb 3 Sn as weil as on niobium [6 ]. The primary and secondary inductors were made witb 50-~J.m-wide lines on 75 IJ.ID centers and bad 25 and 58 turns, respectively. Eacb inductor bad an inside diameter of 3.00 cm and an outside diameter of 3.86 cm. The values of L 1 and L 2 were calculated CJ to be 40 and 210 IJ.H. The superconducting properties of tbe circuit elements made from the sputtered niobium films were obtained by standard dc measurements. The transition temperature was found to be 9.3 K and tbe dc critical current of tbe coils was about 90 mA. The typical resistivity ratio between room temperature and a temperature just above tbe transition temperature was about five . RESULTS AND DISCUSSION Measurements were carried out at 4.2 K witb tbe circuit assembly sbown in Fig. 2, immersed directly in liquid belium. Signals were transmitted into and out of tbe cryostat with tbe help of 50-0 semirigid coaxiallines. The signal source used was an HP 606-A signal generator, and tbe input and output voltages were measured on an oscilloscope witb a 50-0 termination. The overall-frequency response of the filter was obtained by performing point by point measurements. For illustration purposes, a typical curve of tbe output to source signal ratio vs. frequency over the entire frequency range is sbown in Fig. 3. Tbis was taken witb tbe input signal Ievel at A Superconducting RF Notch Filter 247 0.75 .------,-----,---l ----,---;l-----,-- - , - - - - -- , - - r - --TI I ,;' ..~050 -r .. z C> v; "-' ~ 0.25 -' "-' a:: I 1 0 o~-~--~ 4 --~-~~-~,o~-,,~ 2 -~1~ 4 --t~6--~ ~s~ FREQUENCY. MHz Fig. 3. Overall frequency response of filter. 40 mV. Anode exists in the low-frequency range, the output signal rises rapidly with frequency and Ievels oft at higher frequencies when the impedance of the filter circuit becomes greater than the 50-n Ioad resistance. As the frequency increases further, the voltage ratio remains at about 0.5. The notch of the filter appears as a sharp dip of the output voltage. The detailed shape of a notch on an enlarged frequency scale is shown in Fig. 4. From a series of measurements, a plot of the reciprocal of the square of the notch frequency vs. the tuning capacitance was obtained, as shown in Fig. 5. The linear dependence of 1/f~ with C 2 is in accordance with the relation in (3). From this plot the stray capacitance of L 2 and the associated capacitor plates is found to be about 3.2 pF from a straight-line extrapolation. The existence of this capacitance imposes an upper Iimit of 11 .7 MHz on the tuning range. Furthermore, the same results allow the calculation of the actual inductance of the secondary coil, which is found to be 53 J.LH. It should be noted that this value is about four times smaller than the nominal inductance calculated from the geometry. In addition, from an independent experiment, by measuring the resonant frequencies of the tank circuits of a capacitor connected in series with the coils, the ratio of inductance between L 1 and L2 was obtained to be 0.75, and the above observed value of L 2 was confirmed. The 0 ~ -2 ~ -4 ~ ...,_~ · 6 !ii -8 - 10 -12 Fig. 4. Detailed shape of a typical notch. 5.90 5.92 FREQUENCY , 11Hz 5.94 C. S. Pang et td. 0.03 ':'~ 0.02 :z: "' 0 4 CAPACITANCE . pf 8 12 Fig. 5. Dependence of notch frequency on tank circuit capacitance. reduction of inductance in L 2 appears to be the result of magnetic field screening produced by the superconducting plates which are located in proximity to the coil. As shown in (1), the depth is sensitive to the resistance R2 and the coupling coefficient k. By increasing the coupling, the maximum notch depth observed was about 17 dB for the nominal 50-0 loading conditions. Values of k and R2 can be calculated from the measured depth and bandwidth of the notch from (1) and (2). The values of k obtained in this fashion are smaller than theoretically predicted based upon the inductance of L 1 and L 2 , and the separation of the coils. values This result indicates that the presence of superconducting materials near the coils reduces the ftux coupling, and that magnetic shielding has to be taken into account in a realistic calculation of the coupling coefficient. The resistance of the tank circuit obtained in this fashion is given in Fig. 6 as a function of frequency. From this graph, the resistance is found to be proportional to w n with n approximately equal to 1.6 over the entire frequency range . n ... 0 z ~ (/) Cii ~ 5 ... 0 ~ "'::>(/) R er f 1. 57 3 1. 5 L--L..---'--........L.........L.---l......L..L...I....L..J....L..J'--__J 3 5 10 15 20 F REQ UENCY , M Hz Fig. 6. Resistance of secondary coil as a function of frequency at 4.2 K. A Supercondudinc RF Noteh Filter Z49 The maximum power handling capability of the filter is about 4 mW, which corresponds to about 50 mA and 200 V of current and voltage being developed in the tank circuit. This is in reasonable agreement with critical current measurements. In the frequency range of 2 to 12 MHz, the output to source signal ratio is independent of the signal strength within the 4-mW maximum just discussed, which indicates that the resistance does not vary with the current in contrast with the results of Blair et al. [8 ], who find that the rf loss increases with signallevels. Finally, the quality factor Q of the present filter is about 2000 with 50-.0 source and output impedance while the circuit is fully loaded. Theoretical calculations indicate that at temperatures below Tc/2 and frequencies lower than the pair-breakin;, frequency, the intrinsic resistance is proportional to the square of the frequency [ ]. For real materials, however, the resistance is also determined by other material properties [ 1 0-- 16], such as (1) surface roughness and the oxide layer on the superconductor, (2) trapped magnetic vortices and normal domains in the material, (3) magnetic hysteresis effects of the ftux motion, (4) Iosses by magnetic coupling to normal metals, and (5) dielectric loading. Recently, Judish et al. 7 ] measured the surface resistance of a helically loaded Iead cavity at frequencies from 136 to 472 MHz. Extrapolation of their data to 10 MHz and 4.2 K yields a value of the surface resistance of about 10-9 .0, which is within a factor of 20 of the measurements obtained in this study on the niobium film (normalized to the same low-temperature resistivity). Although the dominant loss has not been identified in this case, the strong frequency dependence of the observed resistance rules out the possibility of dielectric loss in the substrate materials. Such a loss depends on the dielectric loss tangent which is only a weak function of frequency in the range of the present measurements 8 ]. Hysteresis loss can also be ruled out since this loss should depend linearly on frequency and a dependence of the loss on signal strengths has not been observed. e e SUMMARY A preliminary investigation has been conducted of a superconducting notch filter for possible application in the 2 to 30 MHz high-frequency (HF) communication band. The circuit was successfully implemented using planar geometry so that closed-cycle refrigeration could be used to cool circuits fabricated from high Tc Nb3Sn or Nb3Ge thin films. In the present design, circuit Q's of about 2000 were obtained with a 50-.0 source and output impedance. Circuit Q's of about 1000 to 2000 are required in ordertoperform fittering of signals in the HF band; the high Q's available with superconducting technology coupled with the possibility of implementing a wide tuning range outperforms conventional tunable notch filters. Conventional HF notch filters have typical circuit Q's of about 75 to 100 at 10 MHz and are typically limited to an octave tuning range. The maximum input power to the filter was found to be about 6 dBm, which enables the superconducting filter to be used to protect receiver front ends from strong HF intederence signals. Measurements indicate the rf critical current is comparable to the dc critical current, thus providing a means for estimating the maximum power handling capability. The undesirable effects of magnetic ftux shielding on L2 has led to an improved design utilizing rutile (Ti0 2 ), a low-loss and high-dielectric-constant material, to reduce the capacitor plate area. Knowledge of the resistance and its frequency dependence can be used to predict the performance of resonators at other frequencies. In this 250 C. S. P11111 et al. preliminary investigation the dominant source of loss has not been uniquely identified, although the results indicate that dielectric or hysteresis Iosses are not dominant. ACKNOWLEDGMENTS The authors would like to thank K. E. Gray for useful conversations and T. R. Werner for experimental assistance. RE FERENCES 1. A. I. Braginski, J. R. Gavaler. G. W. Roland, M. R. Daniel, M. A. Janocko, and A. T. Santhanam, IEEE Trans. Magn. Mag-13:300 (1977). 2. R. E. Howard, C. N. King, R. H. Norton, R. B. Zubeck, T. W. Barbee, and R. H. Hammond, in Advances in Cryogenic Engineering, Vol. 22, Plenum Press, New York (1975), p. 332. 3. R. T. Kampwirth, J. W. Hafstrom, and C. T. Wu, IEEE Trans. Magn. Mag-13:315 (1977); and R. T. Kampwirth, C. T. Wu, and J. W. Hafstrom, in Advances in Cryogenic Engineering, Vol. 24, Plenum Press, New York (1978). p. 465. 4. Cutter/Hammer, Airborne Instrument Lab., Final Report, Contract DA ll-119AMC-02530(4) (Dec. 1966). 5. DOD Tech. Memorandum, Rept. No. S-2111. 633 (July 1977). 6. R. T. Kampwirth, I. Schuller, and C. M. Falco, patent pending (1978). 7. F. W. Grover, Inductance Calculations, Dover Publishing Company, New York (1962). 8. D. G. Blair and W. 0. Hamilton, Rev. Sei. Instrum. 50:279 (1979). 9. D. C. Mattis and J. Bardeen, Phys. Rev. 111:412 (1958). 10. T. A. Buchhold, Cryogenics J:141 (1963). 11. J. L. Zar, J. Appl. Phys. 35:1610 (1964). 12. C. R. Radden and W. H. Hartwig, Phys. Rev. 148:313 (1966). 13. J. M. Pierce, J. Appl. Phys. 44:1342 (1973). 14. S. Giordano, H. Hahn, H. J. Halama, C. Varmazis, and L. Rinderer, J. Appl. Phys. 44:4185 (1973). 15. J. M. Victor and W. H. Hartwig, J. Appl. Phys. 39:2539 (1968). 16. P. Kneisel, 0. Stoltz, and J. Halbritter, IEEE Trans. Magn. Mag-15:21 (1979). 17. J. P. Judish, C. M. Jones, F. K. McGown, and W. T. Milner, Phys. Rev. B 15:4412 (1977). 18. Reference Data for Radio Engineers, ITT, Howard W. Samf Company, New York (1973). DISCUSSION Question by D. Petrac, Jet Propulsion Laboratory: How many turns does the inductor have and what are its inductance values? Answer by author: There are 25 and 58 turns in the primary and secondary coil, respectively, with a mean diameter of about 3.4 cm. The calculated values of inductance are 40 and 210 ~-tH. Question by C. H. Morgan, Brookhaven National Laboratory: Did you consider geometrics other than flat spirals, e.g., coaxial coils? Answer by author: Yes, besides the planar configuration presented here, we have also considered a geometry in which the inductors were fabricated by etching thin-film coating on a hollow dielectric cylinder. Question by R. C. Longsworth, Air Products and Chemicals, Inc.: How was the device cooled during the experimental work? Answer by author: In this measurement the filter assembly was immersed directly into a liquid helium bath. Question by R. C. Longsworth, Air Products and Chemicals, Inc.: What will the closed-cycle refrigerator requirement be in terms of temperature, temperature stability, capacity, vibration, etc.? Answer by author: It should be kept in mind that we are only in the preliminary stages of a feasibility study period. All of the details of filter performance, operation, and construction depend on the results of the study. These have not all been worked out. The goal of the first part of the study is to establish the limitations of filter performance. Wehave not yet measured the performance as a function of temperature. So far, the study suggested that the operating temperature of a Nb3 Sn filter should be kept below about 12 to 13 K and it is expected that performance will be insensitive to temperatures below this point. E-3 EXPERIMENTAL EVALUATION OF A 1-METERSCALE D-SHAPED TEST COIL FABRICATED FROM A 23-METER LENGTH OF INTERNALL Y COOLED, CABLED SUPERCONDUCTOR* M. 0. Hoenig, A. G. Montgomery, and S. J. Waldman Francis Bitter National Magnet Laboratoryt Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCTION Superconducting magnets have generally been cooled by means of pool-boiling helium within a cryostat vessel. The internally cooled, cabled superconductor ICCS, on the other hand, requires only the circulation of helium through its passages. In the 1-m-scale cryogenic facility the test coil is suspended in a thermally shielded vacuum where it is exposed to a background magnetic field, provided by a pair of split solenoids, clamped across its central, straight section (see Fig. 1). Refrigeration and helium ftow are su,pplied by a Model 1400, CTI refrigerator. In past work 3 ] various internally cooled, cabled conductors were subjected to stability tests. Such coils bad the shape of a tightly wound cable, protected from Lorentz forces by tensile winding restraint. The primary objective of the 1-m-scale test facility is its ability to expose straight lengths of unsupported conductor, such as the side of a D-shaped coil, to Lorentz force. e- TEST COIL Figure 2 shows the test coil installed. The coil itself was wound in a single-layer pancake 152 cm talland 71 cm wide using six turns of conductor. The aluminumsheathed conductor was placed in an anodized aluminum extrusion (see Fig. 3) and was cowound, with the extrusion into the coil shape. Five turns of the conductor, one electrical termination, and one hydraulic connection are shown in Fig. 4. CONDUCTOR Figure 5 shows a section of the conductor, which consists of 19 sets of triplets. Each triplet is a twisted set of three strands, each 1.05 mm in diameter. The coil was formed by twisting first 6, then 12 triplets around a central triplet, with each layer * Work supported by the U. S. Department of Energy. t Supported by the National Science Foundation. 251 151 M. 0. Hoenlg, A. G. Montgomery, and S. J. Waldman Fig. 1. D-shaped test coil shown with split-pair solenoid. Fig. 2. D coil, split-pair solenoid and helium canister, supported by 15-K top plate. Experimental Evaluation of a D-Shaped Test Coil 253 Fig. 3. Aluminum-sheathed cable and anodized aluminum extrusion. Legend: 1, Coil space; 2, aluminum tube (12-mm diameter); 3, aluminum extrusion (anodized); 4, fiberglass insulation (epoxy impregnated); 5, insulation overwrap (epoxy impregnated); 6, epoxy filled space. individually transposed. The cable is contained in an aluminum tube with a 0.9-mmthick wall. The helium cross section inside the tube represents 40% of the total cross-sectional area. SUPERCONDUCfOR The superconducting wire was fabricated by MCA. * Each strand has a diameter of 1.05 mm and consists of 83% copper and 17% NBTi in the form of 671 filaments of approximately 8-IJ.m diameter. The same cable and superconductor have been tested previously [ 1] in the form of a 15-cm-diameter test coil. BACKGROUND FIELD COILS A 7-T background field was provided for the test coil by means of a pair of iron-filled split solenoids, fabricated by MCA, and illustrated in Figs. 1 and 2. The solenoids were attached to the test coil in the manner shown in Fig. 6. A plan view of the assembly is shown in Fig. 7. * Magnetic Corporation of America. 254 M. 0. Hoenig, A. G. Montgomery, and S. J. Waldman Fig. 4. Section of coil shown with electrical and hydraulic terminations. Fig. 5. Magnified cross section of conductor, showing aluminum sheath and cabled conductor with NbTi filaments (hexagonal shape). FORCED·FLOW COOLING AND REFRIGERATION A simplified schematic of the forced-flow helium cooling system is shown in Fig. 8. Coldhelium ftow to the test coil is obtained from the cold, high-pressure side of the refrigerator. The gas at 3 to 4 atm pressure is cooled down to coil-operating temperature (4.7 to 5.3 K) by means of heat exhange to 1 atm liquid helium within a canister reservoir shown in Fig. 2. 2SS Experimental Evaluation of a D-Shaped Test Coil Fig. 6. Adapterblock and coil with pulse coil, ready for bolting between split-pair solenoids. 0 AOAPTER _ BLOCK 0 0 0 45" 01A , 15K THERM AL SH ROUO 18" 39" Fig. 7. Plan view of assembly, showing cross section of test coil. Helium enters and leaves the test coil as a single-phase fluid. Mixed with additional bypass coolant, it passes through the thermal shroud, maintaining its temperature close to 15 K. Downstream of the shroud the helium is expanded to 1.2 atm for return to the refrigerator. The refrigerator, with its two standard compressors, has a nominal capacity of 7 5 W. Since this is not adequate to satisfy the total requirements, additional helium is supplied during steady-state operation to maintain liquid Ievel in the helium canister. TEST FACILITY The integrally tubulated stainless steel shroud assembly is illustrated in Fig. 9.1t is cooled to about 15 K by the refrigerator. The shroud provides a 172-cm-tall, M. 0. Hoenig, A. G. Montgomery, and S. J. Waldman 256 WARM RETURN TO COMPRESSOR COLD 300A LEADS :12777d7777ZZZ;6ZZZ7~ 1 [~ VACUU M SPACE ~==)VACUUM CHAMB ER = --=--"' h Figo SoSimplified schematic of forced-flow helium cooling systemo ROOM TEMP. TOP PLATE 78 K o LNz COOLED PL ATE 15 K o HELIUM COOLED PLA TE 45°0 0 D SUPPORT RING _____... TEST 68" VOLUME 44"0 D COI L SUPPORT RING 15K WALL HELIUM COOLE D CY LI NDER - - - -- HELIUM COOLED BOTTOM PLATE Fig. 9. Elevation view of 15-K, evacuated test volumeo Elqteriolental Evaluation of a D-Shaped Test CoD 157 114-cm-diameter, 1. 75-m3 working volume within its optically tight enclosure. The shroud is brought up vertically into position, is bolted to its 15 K top plate, and thus surrounds the test coil and its apparatus. The 15-K structure is suspended from a liquid-nitrogen-cooled stainless steel plate and support ring, itself suspended from the top, room temperature support plate of the vacuum chamber. INSTRUMENTATION The dewar and D coil were instrumented with 20 carbon resistors to indicate steady-state temperatures. Calibrated carbon glass resistors were attached to the liquid helium inlet and outlet of the D coil, as weil as the orifice meter. Static helium pressures in the system were measured with room temperature transducers connected by means of capillary tubes. Helium pressure transients were measured with a cryogenic strain gage pressure transducer. This 0-150 psig transducer was placed upstream of the D coil as shown in Fig. 8. A voltage-tap pair was connected to each of six turns of the D coil. On each conductor turn, pair connections were placed on opposite sides of the high magnetic field region and one of the Iead wires was brought through the high-field region to meet the other in order to minimize inductive pickup in the voltage tap pairs. A two-layer pulse coil was wound onto the D coil. This pulse winding is represented in Fig. 6. A capacitor bank charged from 80 to 500 V could be discharged into the coil to produce a 0.3- to 2.2-T pulse with a 12-ms width. Although a significant fraction of the field may be screened out by the 0.9-mm-thick aluminum sheath, the penetrating magnetic field pulse induces currents within the cable itself within the conduit. A pulse coil calibration was performed on a separate length of the conductor in zero field, allowing for the correlation between a capacitor discharge voltage and the deposition of heat energy in the cable. A capacitor discharge voltage of 300 V was tentatively correlated to the deposition of 270 mJ I cm3 of cable metal. A cryogenic pressure transducer was used to examine the thermodynamic state of the pressurized helium, approximately 5 ms, after pulse discharge. The capacitor voltage discharge was thus related to the indicated internal energy change in helium. Initial helium temperature and pressure within the cable space were 4.2 K and 3 atm, respectively. Helium ftow through the D coil was measured by means of a cryogenic orifice plate ftow meter. The ftow meterwas located downstream of the D coil as shown in Fig. 8. The ftow meter was calibrated by passing helium through the cryogenic ftow meter and a room temperature volumetric gas meter at a constant-mass ftow rate. TEST COIL OPERATION The test coil has a critical current rating of 6800 A at 5.0 K and 7 T. lt can be subjected to pulsed eddy-current heating. The following procedures were used in its operation. Pumpdown and Cooldown The 4-ft-diameter by 8-ft-tall vacuum chamber was evacuated by means of a 6-in diffusionpump to about 2 X 10-5 torr within about 6 hr. The heliumsystemwas purged and connected to the refrigerator, with both systems at room temperature. Initial cooldown to 100 K took about 24 hr and was carried out by the refrigeration system using its liquid-nitrogen-cooled heat exchanger. Subsequent cooldown to 10 K took another 12 hr using the full output of the two expansion 258 M. 0. Hoe.U., A. G. MHf~GmerY, IUid S. J. Waldaum engine Model1400 CTI refrigerator. As noted from the schematic (Fig. 8), helium ftow can be regulated to cool down the cabled conductor and thermal shroud in series, followed by the D coil's current Ieads and the split-pair solenoid with its Ieads, in parallel. Gas ftow is then returned to the refrigerator, partly through the warm output of the Ieads. The capacity of the refrigerator is adequate to bring the D coil temperature to approximately 10 K and the split pair solenoid and thermal shroud temperature to 20K. Preparadon for Test Operadon Final cooldown of the system can be carried out in less than 1 hr. Liquidhelium is transferred to the canister, located within the thermal shroud, and connected to feed the split pair solenoids by gravity. The system is ready for test Operation once the following requirements have been satisfied: 1. Split-pair solenoids must be fully ftooded with liquid helium from the canister. Once the 300-A Ieads to the solenoids are adequately cooled, the solenoids can be charged up to the current Ievel corresponding to the desired field. 2. The D coil 7000-A current Ieads must be adequately cooled. 3. The D coil inlet and outlet temperatures must be at or below 5 K and steady. 4. Helium pressure and ftow rate through the test coil must be established. Test Operadons Steady State. Steady-state operations were performed at 6 and 7 T for extensive periods of time with a dc current of 6000 A. No quench of the test coil was observed. Critical Cu"ent Tests. The coil was brought up to its critical current in a background field of 7 T and a helium temperature of 5.0 Kin the cable. Quench occurred at 6720 A corresponding to a current density of 817 A/mm 2 (NbTi), matehing the short sample current density of the superconductor at 5 K and 7 T. As observed in the schematic (Fig. 8), helium can only expand out of the coil through its downstream connection. This was arranged in this manner to measure pressure buildup (on the upstream end of the conductor) as the result of a quench. Starting with an initial pressure of 2.65 atm, a peak pressure of 6.6 atm was noted 1 s after the appearance of voltage, indicating a normality. Cu"ent Sweep and Reversal Tests. Several current sweep tests were performed to see if wire motion in the cabled conductor in its straight section would generate sufficient heat to cause a quench. With the background field at 6 T, the current, I, was ramped up to 5500 A in 25() ms. Peak dl/ dt was 40,000 A/s at zero time. Though the temperature of efB.uent helium was slightly elevated, no quench occurred. Steady-state conditions were established at 5500 A. In a similar test the conductor quenched in an attempt to reach 6000 A in 250 ms. Peak dl/ dt, in this case, was 44,000A/s. The direction of the current was reversed several times in order to see the effect of possible bunching of cable strands. No effect was observed. Pulse-Heating Tests. Table I shows the results of numerous pulse-heating tests. With a given current, field, helium ftow rate, and helium pressure established, the conductor was heated with a pulse coil. The data showing maximum stability have been plotted in Fig. 10, in terms of approximate thermal energy input (mJ/cm3 ) per unit volume of wire vs. 1/10 , the normalized current. Experimental Evaluation of a D-Shaped Test CoD ...E . u ~ "'...J <[ u 259 0 v N 0 0 N (/) >(!) Q: "'z "' 1"' 0 !!! 0 ~ <[ :::0 x 0 Q: ll. ll. <[ 0 CD ... 0 Fig. 10. Stability curve; pulsed energy vs. normalized current. 02 04 06 08 10 NORMALIZED CURRENT, 1/lo CONCLUSIONS AND DISCUSSION Steady-state Operations of the D coil were performed at 89% of its quench current of 6702 A (7 T) at a temperature of 5 K. The critical current of the coil was found to match its short sample quench current at 7 T and 5 K. Rapid (250-ms) current sweep tests demonstrated that the conductor current could be brought up from 0 to 5500 A without quench, with a peak, dl/ dt, of 40,000 A/s. This ramp stability is viewed as evidence that Lorentz force generated wire movement, if any, produced insignificant conductor heating at this operating Ievel. Pulse heating tests indicated recovery from a range of heating pulses at both 6 and 7 T. These results are plotted in Fig. 10. The coil was Operated at a reduced helium velocity (Table I) with no change in stability e]. While normal Operations took place at a helium pressure of 3 atm, a test run was made at 4 atm in order to determine the effect of increased pressure on performance. No effect was found. In no instance were extreme pressures observed upon nonrecovery. It is believed that because the high magnetic field is confined to about 4% of the conductor under test, the normalcy does not propagate into the low field (96%) of the D coil. Table I. Stability Test Data Field, T Temperature,* K Pressure,* atm 7 7 7 6 6 6 6 5 5 5 5 5 5 5 3 3 3 3 3 3 * Helium properties. 3 Velocity,* cm/s 18 18 6.6 18 18 18 18 A Normalized current Pulse time, ms Pulse voltage, V 2900 4700 4700 5600 6500 6000 2000 0.43 0.70 0.70 0.83 0.97 0.89 0.30 12 12 12 12 12 12 12 180 80 80 125 80 105 280 Current, M. 0. Hoenig, A. G. Moatgomery, lllld S. J. Waldman Testoperations were performed at conductor temperatures of 4.7 to 5.3 K. Since the NbTi critical temperature Tc at 7 T is only 6.1 K, the thermal margin was very limited. Relatively high stable energy inputs (Fig. 10) would thus tend to indicate rapid recovery and very short normal and current-sharing periods. REFERENCES 1. M. 0. Hoenig and A. G. Montgomery, in Proc. 7th Symp. Engineering Problems of Fusion Research, Vol. I, IEEE Science Center, Piscataway, New Jersey, (1977), p. 780. 2. Y. lwasa, M. 0. Hoenig, and D. B. Montgomery, IEEE Trans. Magn. Mag-13:678 (1977). 3. M. 0. Hoenig, J. W. Lue, and D. B. Montgomery, in Proc. 6th Intern. Conference on Magnet Technology, ALFA, Bratislava, Czechoslavakia (1978), p. 1021. E-4 PERFORMANCE OF GAS-FILLED THERMAL SWITCHES J. Yamamoto Osaka University Suita, Osaka, Japan INTRODUCTION A thermal switch has been developed for precooling a superconducting magnet with a cryocooler C· 2 ]. The switch is essentially a heat pipe [3 ], which provides a high heat transfer rate when the working gas is being liquefied and acts as an insulator when the dewar for the magnet is filled with liquid helium. Configuration and performance of the switch are described herein. CONSTRUCTION The thermal switch consists of a gas-filled stainless steel tube (262 mm lang, 15 mm in diameter and 0.3 mm in thickness), with cone-shaped copper ends as illustrated in Fig. 1. The encapsulated gas begins to liquefy when the temperature at the upper end reaches the condensation temperature of the gas. As the gas condenses on the upper cone, droplets form and fall off the apex onto the lower cone, where they evaporate. This process of condensation and evaporation in the tube transfers heat from the lower end of the tube to the upper end. While this process is occurring, the switch is said to be in the "on" condition. The temperature range or band during the "on" state of operation is wider when the charging pressure of the gas is increased. During a typical operation the thermal switch is filled with either nitrogen or hydrogen at 1.3 MPa (13 atm) at room temperature. When either end of the thermal switch is below the triple point of the gas (13.9 K for hydrogen and 63.2 K for nitrogen), the switch automatically reverts to the "off" condition owing to the solidification of the fluid. PERFORMANCE Figure 2 shows the relation between the gas pressure in the tube and the temperature of the upper end (Tu) and the lower end (T1) for the nitrogen-filled switch when a small copper piece (about 50 g) is attached on the lower end, and the upper end is cooled by a cryocooler. The saturated vapor pressure curve for nitrogen is also shown in the figure as a dotted line. A dashed line shows the change of the inside pressure. As the temperature of the upper end is lowered, the inside pressure decreases and T 1 slowly decreases, because the switch is "off". When Tu matches the saturation 261 J. Yamamoto 262 gas inlet tube ..J...i-- - - - - - - copper cone - - - - - stainless steel tube (15 mm od, 0.3 mm thicknessl E E N !El copper cone Fig. 1. Configuration of thermal switch. Stainless steel tube is capped by copper end pieces. After filling with working gas at 1.3 MPa, the inlet tube is sealed off. 1. ' I l II! 0.. o. i ::i'; N2 >' : a. 1.0 ""' I! .'"I. -· I ~i 0 .5 T.K Fig. 2. Pressure vs. temperature curve for a nitrogen-filled switch during the precooling stage. Dashed line shows an isothermal change. Performance of Gas-Filled Thermal Switches 263 o. a .---~--.----..----.--.------. CO 0.6 Q. ::E 0.4 0.2 Fig. 3. Pressure vs. temperature curve for a H 2 -filled switch during the precooling stage. T. K temperature of the enclosed gas, the nitrogen gas starts to liquefy, thereby activating the thermal switch. Tu follows along the saturated vapor-liquid line of Fig. 2 when the switch is in the "on" state. Finally T1 is coincident with Tu on the saturated vapor-liquid line. The same situation is observed for the hydrogen switch as shown in Fig. 3; however, the saturated vapor pressure line corresponds to lower temperatures. The thermal switches which were used to provide the data in Figs. 2 and 3 were not sealed off but bad an extension tube for measuring pressures. The volume of the extension tube is about 5% of the total volume of the switch itself. The measured heat transfer rates of the switches as a function of Tu are shown in Fig. 4. To obtain these heat transfer rates, a small electrical heater was mounted on the lower end of the switch. The heat transfer rate was determined by dividing the applied electrical heat at the lower end by the temperature difference of the two ends. Hz swltch 100 r", i I I I ' I I r I I \ lI N2 switch . I 'I + I I I I I \ Fig. 4. Heat transfer rate of the nitrogen and hydrogen gas-filled switches. Abscissa depicts temperature of the upper end of switch. .... 0'~-~-~~----,~~-~-.h,; len"4)llrtllure , K J. Yamamoto 15 3:1.0 i ~ 2 ~ 05 0 2 6T. K 3 4 Fig. 5. Heat transfer performance of the hydrogen switch in the "on" condition. Abscissa represents the temperature difference, T1 - Tw. The heat transfer rate was measured under conditions of low applied heat so that it would be independent of the temperature diflerence. This provides a close approach to the isothermal conditions noted in Figs. 2 and 3. The nitrogen switch provided a heat transferrate of 3 x 10- 1 W /K during the "on" condition (T.. range of 63 to 88 K). The hydrogen switch showed a very similar heat transferrate although the T.. in the "on" condition moves to a lower temperature (16 to 20 K). The heat transfer rates shown in the figure include the contribution of the stainless steel tube. This is estimated as 3.6 x 10-4 W /K and 4.3 x 10-s W /K at 60 K and 10 K, respectively. With T.. lower than the triple point of the filling gas and temperatures below the "on" condition, the heat transfer rate is about 5 x 10-4 W /K; when compared to that estimated for the stainless tube alone, one can see that the heat transfer is primarily determined by the tube conduction. At temperatures above the "on" condition, the heat transferrate shown in the figure (about 5 x 10-3 W /K) is a result of the natural convection of the filled gas. Figure 4 shows the heat ftow of the hydrogen switch during the "on" condition. The heat ftow is proportional to the temperature diflerence between the two end cones up to 1.1 W. When additional electrical heat is applied at the lower end of the switch, a saturation phenomenon occurs which may be caused by a limited condensation rate at the surface of the upper cone. CONCLUSION Application of the thermal switch to a real cryostat with a cryocooler results in an eflective "on" region which is wider than that shown in Fig. 4 because a large temperature diflerence is caused in the precooling stage (see Figs. 2 and 3). The "on" temperature range can be incremented up or down and its duration increased or decreased by changing gas composition and pressure. The perfonnance of the switch is very attractive for many applications in cryogenics. The application. of the switch for grecooling a superconducting magnet has been successfully examined elsewhere [ ]. ACKNOWLEDGMENTS The author wishes to acknowledge the services of M. Yanai of Osaka Oxygen Industries, Ltd. for designing the thermal switch and the helpful discussions ofT. Shigi and Y. Inuishi during this work. Performance of Gu-Filled Thermal Switdles 265 RE FE RENCES 1. J. Y amamoto and T. Shigi, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Press, Guildford, England (1978), p. 593. 2. J. Yamamoto and M. Yanai, Rev. Sei. Instrum. 50:1382 (1979). 3. R. P. Bywaters and R. A. Griffin, Cryogenics 13:344 (1973). DISCUSSION Question by E. Lady, University of Michigan: Is the heat transfer of the thermal switch limited to the vertical direction? Answer by the author: Yes, the droplet of the liquefied gas transports heat in the thermal switch. F-1 TRANSIENT COOLING OF A FAULTWORTHY SUPERCONDUCDNG ELECTRIC GENERATOR* J. A. Schwoerer and J. L. Smith, Jr. Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCDON In order for superconducting generatorstobe of use in power stations, they must be made faultworthy, or able to withstand the effects of a short circuit on the power lines. The work described herein is part of an MIT-DOE program to design a 2000-MVA generator, the size of commercial interest, and to build and test a 10-MVA prototype. Designs of faultworthy superconducting generators to date have stationary, normally conducting armatures, rotating superconducting field windings, and several rotating electromagnetic shields between the field winding and the armature [ 1- 3 ]. The MITdesign has two electromagnetic shields and is unique in that these are structurally part of the rotor and operate in the steady state at temparatures in the range of 4 to 6 K. A system of cold shields, as opposed to having at least the outermost shield self-supporting and at room temperature, simplifies the structural problern and has other advantages, but it makes the cooling problern much more difficult [ 1]. Fora representative 2000-MVA design, subjected to a severe fault used for design purposes, the time-averaged total losses in the shields are 4 MW during the first 0.1 s and thereafter drop by an order of magnitude and decay exponentially. The transient cooling scheme must prevent this heat input from increasing the temperature of the superconductor to the level where it will go normal. The principal feature of this transient cooling scheme is a self-pumping helium ßow system for removing the heat input to the shields. This works in combination with thermal isolation layers which protect the superconductor during the short time that the temperature of the shields is high. The transient cooling scheme is described and analyze<\ below. This analysis has been used to show that the cooling concept is viable for both 10- and 2000-MVA sizes. COOLING SCHEME DESCRIPTION Figure 1 is a preliminary design layout of the 10-MVA generator. The outer of the two electromagnetic shields is the damper. It is a copper winding connected to an extemal resistor and is designed both to damp out rotor swings and to shield the superconductor from unsteady fields. The inner of the electromagnetic shields is the * Work supported by U. S. DOE under Contract No. EX-76-A-01-2295, Task Order No. 11. 266 SUPPORT TUB E I He COUPL I NGS VACUU M AI R G A P T0 10 C M STA T OR BOR E SEAL O UT ER I Fig. 1. MIT 10-MVA preliminary design Iayout. HOOP SUPPO RT STAND - OFF R ~GS S O L I D SHIEL D TO POWER COUPLI N<"; T HERMAL ~ i ftl ; "~ :1. ~ ii' l"l ! Ii s= 1 ~ a-~ ri .,2. Jf 2. ~ a ~· ~ J. A. Schwoerer and J. L. Smith, Jr. SH I ELO HE LI UM RESERVOIR FLOOOEO SHI ELO CHECK VALVE Fig. 2. Schematic of shield cooling system. solid shield. It is a solid copper cylinder designed to intercept the unsteady magnetic fields that leak through the damper. During a fault, approximately 95% of the rotor heat input occurs in the damper in the form of 12 R losses, approximately 5% occurs in the solid shield as 1 2 R losses, and « 1% occurs in the field winding as magnetic hysteresis Iosses in the superconductor and 12 R Iosses in the copper matrix Cl. The self-pumping helium ftow systems for cooling the solid shield and damper are separate and identical in the generalized design analyzed here. The system for either shield is diagrammed in Fig. 2. It consists of a reservoir near the end of the rotor, helium passages in the shield connected to the reservoir by a line containing a check valve, and a vent line leading from the end of the shield opposite the reservoir to the rotor centerline and then out to the atmosphere. The ftow is throttled by a nozzle at the end of the vent line. Recently, consideration has been given to a modified design in which the helium cooling systems for the damper and solid shield are interconnected and there is a single check valve that is near the centerline between the field-winding reservoir and one of the shield reservoirs [ 1]. The other parts of the transient cooling system are the thermal isolation layers and the cooling system for the field winding. The 10-MVA design, shown in Fig. 1, has stainless-steel-epoxy thermal standoft rings between the damper and the solid shield. The stainless steel outer support tube doubles as a thermal isolation layer between the solid shield and the field winding. The field winding is ftooded with helium and connected to the field-winding reservoir by radial passages containing copper isothermalizing bars. The operation of the cooling system is as follows. Because of the centrifugal acceleration, the pressure of the helium in the shields and the field winding is always above the critical pressure ; thus, boiling does not occur. In each shield the transient heat input causes the pressure to rise, closing the check valve and producing ftow out the vent line. The resultant expansion produces a cooling effect and counteracts the pressure rise. When the pressure in the shield again becomes equal to the pressure at the bottom of the shield reservoir, the check valve opens. This first period of the operation is termed the "blowdown phase." The cold helium that enters the shield from the reservoir partly mixes with and partly displaces the helium in the shield and produces further cooling. The ftow is driven by the thermosiphon effect: the fluid in the vent line is less dense than that in the reservoir. Thus the second period of operation is designated as the "chimney phase." Oscillations in the shield pressure similar to those observed in filling a helium dewar may occur. The resulting shield temperature history is, roughly speaking, a sharp rise followed by an exponential return to the steady-state Ievel. The thermal isolation layers are designed so that the time it takes for the thermal wave to propagate across the layers is long compared to the time it takes for the self-pumping systems to recool the shields. Transient Cooling of a Fanltwortby Snperoondncting Electric Generator 269 ANALYSIS The analysis described here deals primarily with the self-pumping shield cooling system shown in Fig. 2. It has been assumed that the heat capacity of the solid parts of the shield is negligible such that the electrical heat generation is nearly instantaneously transferred to the helium. For cooling channel geometries that have been considered, there is significant enhancement of the heat transfer by the secondary flow driven by natural convection, and the convective heat transfer coefficient has been calculated tobe approximately 0.5 W /cm 2 K [4 ]. While the configuration of the helium passages in the shields has not been finalized, one can conceive of designs, with conductors and helium in direct contact and a large amount of surface area, where the assumption is accurate. In designs to date, the heat transfer between the shields and the adjacent thermal isolation layers has been neglected in the analysis of the self-pumping shield cooling system. This has led to conservative results since either the heat transfer with the thermal isolation layers has been small compared to the electrical heat generation in the shields, or, at a given time, the net heat transfer has been out of the shields. This might not be true for the solid shield in some designs. In such a case, an iterative procedure would be followed in which the heat input to the shield would be adjusted according to the calculated heat transfer with adjacent thermal isolation layers. The design is such that changes in state at the entrance and exit of the shield may be neglected. This is reasonable since the cross-sectional area of the line between the reservoir and the shield and that of the vent line are approximately the same as the combined cross-sectional area of the cooling passages in the shields. Finally, it has been verified by calculation that the kinetic energy resulting from the fluid velocity in the direction of ftow can be neglected, except at the vent line exit nozzle, throughout the transient. For the blowdown phase, when the check valve is closed, it is assumed that the state of the helium in the shield is uniform. This is accurate because heating is uniform, the pressure drops due to viscous and inertial effects are very small, and the axial thermal conductivity of the shield is high. Applying conservation of mass and conservation of energy to an annular control volume around the shield, one obtains dp. dt and V _1_( () _m/•) Ps du, = Vp, dt (1) (2) For the chimney phase, when the check valve is open, two limiting-case models are considered: an "instant-mixing" model and a "no-mixing" model. In the instant-mixing model the state of the helium in the shield is assumed to be uniform. The high axial thermal conductivity of the shield and the fact that the natural convection flows tend to be very strong are factors which breakdown axial variations of the states in the shield. In the no-mixing model it is assumed that the helium that enters the shield from the reservoir does not mix or undergo heat transfer with the helium already in the shield. In both models the reservoir is assumed to be infinite and the shield pressure to be uniform and constant throughout the chimney phase. Recent work has studied the effects of depleting the reservoir [ 1 ]. Also, it 270 J. A. Schwoerer and J. L. Smith, Jr. appears that mixing effects may cause pressure oscillations under some circumstances. For certain combinations of helium states in the shield and reservoir, the instant-mixing model produces a negative ftow into the shield which violates the equations of mixing. This appears to be associated with the fact that when helium of the two states in question mix at constant volume, the pressure drops. However, this problern has not been encountered in recent designs. lt is possible that pressure oscillations will occur only for those cases where the instant-mixing model gives this negative ftow at the inlet. For the instant-mixing model, the conservation of mass and conservation of energy are again applied to an annular control volume around the shield. Changes in u. and p. are related by the constant pressure assumption according to the thermodynamic property (aujap)p,. Combining these results gives 6- mv(h. - du. dt V{p. + [(u.- h,) h,)j(aujap)pJ} (3) and . m, = . mv V + (aujap)p, du. dt (4) For the no-mixing model, the helium in the shield is divided into a number of control mass elements. Applying conservation of energy to an arbitrary control mass results in dhCM 6 dt VpcM (5) The history of mass ftow rate in the vent line determines the time span for which a given control mass element is in the shield. The analysis of the helium ftow circuit is completed by relating mv to the state at the shield exit. In designs to date it has been reasonable to assume that the ftow in the vent line is quasisteady. lt has also been assumed that the ftow is isentropic throughout the vent line and that the helium in the nozzle behaves like an ideal gas. The model is approximate since the ftow may be two phase in the nozzle and helium departs from the ideal gas model for the range of states of interest. For designs to date, the ftow in the nozzle is choked during most of the transient. The equations that describe the self-pumping shield cooling system were solved by computer. For the blowdown phase and for the instant-mixing model of the chimney phase, an Adams-method subroutine developed by Shampine and Gordon was used [5 ]. For the no-mixing model the Runge-Kutta method was simpler for keeping track of the different control mass elements. The data for helium properties needed to integrate the equations were supplied by a subroutine package from the National Bureau of Standards ft The heat transfer across the thermal isolation layers is an unsteady conduction problern with variable-temperature boundary conditions. The problern is nonlinear because the thermal conductivity and heat capacity of stainless steel are strong functions of temperature. The problern was solved by computer using the explicit Euter method. The factor that determines whether the transient cooling system is effective is clearly whether the temperature rise of the superconductor is acceptable. There is electromagnetic heating in the field winding in addition to the conduction across the Transient Cooling of a Faultworthy Superconducting Eledric Generator 271 inner thermal isolation layer. A conservative estimate of the temperature rise of the field winding is that which results if all of the heat input in excess of the steady-state Ievel is stored in the heat capacity of the helium contained in the winding. RESULTS The operation of the self-pumping shield cooling system is demonstrated by the results for the damper of a representative 2000-MVA design. One of the design considerations is matehing the volume of helium in the shield and the minimum area of the nozzle in the vent line to the heat input and shield radius so that performance is optimized. In the present case the damper radius is 82 cm, the helium volume in the damper is 7.9 x 104 cm3 , and the nozzle area is 10 cm 2 . The helium state in the reservoir near the rotor centerline is 1atm saturated liquid. The heat generation during a three-phase fault on the generator terminals Iasting 0.1 s is a constant rate of 3.9 MW for 0.1 s followed by a decayinr, exponential starting at a Ievel fö that of the initial rate with a time constant of 2 s [ ]. This is nearly the warst possible fault that could occur. During the first 0.1 s of the transient, the calculated state of the helium in the shield moves from state A to state Bon Fig. 3. When the rate of heating decreases, the state at first follows path BC on Fig. 3. After the check valve opens, the states in the shield fall on curve AD. Point D corresponds to the maximum temperature at the shield exit predicted by the no-mixing model. The calculated pressure, temperature, and density of the helium in the shield during the blowdown phase are plotted vs. time in Fig. 4. The shield temperature predictions of the two chimney phase models are plotted in Fig. 5. Both the exit temperature and the mixed mean temperature calculated by the no-mixing model are given. The results for density of helium in the shield during the chimney phase are plotted in Fig. 6. The total use of helium from the reservoir is calculated to be 60 kg by the no-mixing model and 80 kg by the instant-mixing model. The results for the operation of the other components of the transient cooling system for the 2000-MVA design are briefty given below. For the solid shield, the Fig. 3. Locus of transient helium states-2000-MVA damper. 4 .0 s, J/gm/K 6 .0 6 .0 J. A. Schwoerer and J. L. Smitb, Jr. 0 .20 100 ...."" E 18 "'0. E " 0 0.. 8 90 0 .16 0 .05 0 .10 t , sec 0.15 Fig. 4. Shield properties during blowdown phase-2000-MVA damper. radius is 68 cm, the helium volume is 4.4 x 104 cm3 , and the nozzle area is again 10 cm2 • The history of heat generation is of the same form as for the damper and has an initiallevel of 0.14 MW. The temperature reached by isentropic compression from the state at the reservoir centerline is 5.3 K. The maximum temperature occurs during the chimney phase and is 6.5 Kat 2 s according to the instant-mixing model and 7.2 K at 4 s according to the no-mixing model for the state at the shield exit. The length of the transient is 16 s according to the instant-mixing model and 8 s by the no-mixing model. The estimates of total helium used for cooling the solid shield are 14 kg (instant mixing) and 10 kg (no mixing). The distance between the solid shield and the damper is set by electrical rather than thermal considerations, and the solid shield is almost 100% isolated from the heat input to the damper. The heat transfer to the field winding has a steady-state level of 350 W if one assumes that the steady-state temperature of the solid shield is reached by isentropic compression from the centerline state in the reservoir while the field winding, cooled by copper bars, is at the same temperature as the reservoir centerline. The peak transient heat transfer rate to the field winding is 700 W at 7 s. 18 30 t, StC 40 Fig. 5. Shield temperature predictions during chimney phase-2000-MVA damper. Transient Cooling of a Faultworthy Superconducting Eledric Generator 273 0 . 18 u ....u ~ <>.. 0 .16 0 . 14 ..__,__..___"L__"L__ _.__ 20 10 0 _.__~---'----=' 30 40 t, 'ec Fig. 6. Shield helium density predictions during chimney phase-2000-MVA damper. The total heat input in excess of the steady-state Ievel is 4500 J and occurs over a period of about 50s. This heat input is large compared to the anticipated electrical heat generation in the field winding. If all of the heat were stored in the heat capacity of the field winding, the temperature rise would be about 0.1 K, which is acceptable. More detailed results for the 2000-MVA design and results for the 10-MVA design are given elsewhere [1 ' 7 ]. The cooling system concept has been found to be viable for both generator sizes. SUMMARY A transient cooling system has been described and analysed for a fault-worthy superconducting generator having two electromagnetic shields mounted on the rotor and operating at cryogenic temperature. The results of applying the analysis to a representative 2000-MVA generator design have been briefly described. Detailed results given elsewhere show that the design concept is viable for both 10- and 2000-MVA sizes. A refined version of the cooling system design presented here will be used in the MIT 10-MVA superconducting generator. NOTATION h = specific enthalpy m, = mass ftow rate from reservoir to shield m. = mass ftow rate in vent line P = pressure (} = total rate of heat generation in the shield t = time since start of transient u = specific internal energy V = shield void volume p = density Subscripts s = shield state (assumed uniform) r = state at bottom of reservoir CM = state of control mass 174 J. A. Schwoerer and J. L. Smith, Jr. REFERENCES 1. Cryogenic Engineering Labaratory and Electric Power Systems Engineering Laboratory, 2. 3. 4. 5. 6. 7. Massachusetts Institute of Technology, "Demonstration of an Advanced Superconducting Generator," U. S. DOE InterimReportE(49-18)-2295, Task Order No. 11, IR 1, Sec. I.B.1; IR4,Sec. II.A.3; IR 7, Sec. I.D, Sec. IV.D; IR 8 Sec. IV.D.2. J. H. Parkerand R. A. Towne, "Superconducting Generator Design," EPRI Final Report EL-557 (1977). M. J. Jefferies and P. A. Rios, "Superconducting Generator Design," EPRI Final Report EL-663 (1978). R. G. Scurlock and G. K. Thornton, Intern. J. Heat Mass Transfer 20:31 (1977). L. F. Shampine and M. K. Gordon, Computer Solution of Ordinary Differential Equations, W. H. Freeman, San Francisco (1975). R. D. McCarty, NBS Tech. Note 631 (1972). J. A. Schwoerer, M.S. thesis, Dept. of Mechanical Engineering, Massachusetts Institute of Technology, Cambridge, Massachusetts (1978). F-2 EXPERIMENTAL SIMULATION OF A CRYOGENIC SYSTEM FOR A LARGE SUPERCONDUCTING ROTOR* L. Sobel, J. L. Smith, Jr., and F. Rumore Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCTION The high-tip speed experiment described in this study involves a cryogenic rotor designed to test an advanced cooling system for the rotor of a superconducting generator. The experiment does not include an actual superconductor field winding but does include a helium chamber typical of the winding space in an actual superconducting rotor. The cooling system performance of the small diameter experimental rotor simulates that of a large-diameter synchronaus rotor by operating at a proportionately higher rotational speed. The experimental rotor has a diameter of about 0.20 m and is designed to operate at 15,000 rpm. This gives a tip velocity of 157 m/s corresponding to that of a 3600 rpm synchronaus rotor with a diameter of 0.84 m. However, all experimentation thus far has been at rotor speeds not exceeding 8000 rpm (tip speed of 83.0 m/s). The major aspects of the rotor's mechanical design have been developed and presented by Tepper [ 1]. A cross section of the rotor is shown in Fig. 1. The design of the cooling system is based upon recommendations presented in a study by Bejan [2 ]. A description of the cooling system giving details of its operation and design has been presented by Tepper [3 ]. A schematic of the complete rotor test system excluding the data acquisition system is shown in Fig. 2. The technical and theoretical achievements of the experimental program are as follows: (1) The thermal and ftow performance of the torque tube has been measured at 8000 rpm; (2) the reliability of the internal instrumentation and the data acquisition hardware and software systems have been established; and (3) the theoretical analysis of boil-off in the rotating frame has been verified experimentally. The cryogenic experiments and analysis are discussed in this paper. The digital data acquisition system with electro-optical coupling to the rotor and all aspects of instrumentation are discussed in other publications [4-6]. Two major problems with regard to rotor operation have also emerged from experimentation. A significant heat leak axially down the centerline of the rotor creates ftow instabilities under certain conditions of rotor operation. The ftow * Work supported by U. S. DOE under Contract No. EX-76-A-01-2295, Task Order No. 11. 275 276 L. Sobel, J. L. Smith, Jr., and F. Rumore (.p\_ TEMPERATURE ~ TRANSOUCER ®-- PR ESSURE TRANSOUCER TOROUE TUBE ®-HEATER 'J _; RADIATION SHIELDS Fig. 1. Cross section of high tip speed rotor. instabilities and some practical steps to avoid instabilities are also discussed herein. Transient "dryout" testing indicates that the geometry of the winding spacereseryoir in this rotor may be inhibiting centrifugal convection of heat from the rotor periphery to the centerline. This aspect of rotor operation is discussed in more detail elsewhere n. Experimental Simulation of a Cryogenic System for a Superconducting Rotor 277 Iranster tube (!) ollen-brodley (!) t hermocouple 0 pressure oooe 0 pressure tronsducer ~moss flowmeter -(-check volve [K]moss flow regulotor Ct:] pressure 0 heot exchonger - ttwotlle J ..!... rotometer rel1ef volve regulotor Fig. 2. Flow diagram of high tip speed experiment. TORQUE TUBE The torque tube is the critical component of the thermal isolation system of a superconducting synchronous generator. The torque tube thermally separates the cryospaces of the rotor from the room temperature environment and transmits generator torque to the cold windings on the rotor. Figure 3 shows the torque tube design used in this rotor. The torque tube is composed of a Micarta tube sandwiched between two thin-wall tubes of 304 stainless steel. The inner and outer tubes carry the structuralloads and the Micarta tube has a series of axial passages and circumferential grooves which serve as passages for the flow of helium vapor which cools the torque tube and thus reduces the conduction heat leak into the cold spaces of the rotor. The thermal performance of the torque tube under various operating conditions is evaluated from the following experimental results: (1) the boil-off rate of liquid helium, (2) the torque tube temperature profile, and (3) the ternperature gradient at the cold end of the torque tube. The flow performance is determined by measuring the torque tube pressure drop as a function of mass flow rate . Torque tube temperature profiles for varying torque tube mass flow rates at a rotational speed of 8000 rpm are shown in Fig. 4. These profiles were made during "dryout" tests discussed later. Thesetests were run with constant torque tube mass ALL EN- BRAOLEY RADIATION THE Rt.IOCOUPLE STAINLESS o,__.____, .5 1 cm Fig. 3. Partial torque tube cross section showing placement of cold end Allen-Bradley resistors. Dimensions: a = 0.81 cm; b = 5.71 cm; c = 0.605 cm; d = 0.203 cm. 278 L. Sobel, J. L. Smith, Jr., and F. Rumore 0 0 "' ,.. 0 N Wo 0::"' =>- I- a:O a::~ w Q_o ::t:N w- o- 0 "' 0 <D 0 "' Fig. 4. Torque tube temperature profiles. Data points are connected for visual aid only. flows and Iasted from 20 to 60 min each. At low mass flows the torque tube time constant was greater than the duration of the test. However, a reasonable identification of the steady-state torque tubeprofileswas possible. The points plotted in Fig. 4 are the temperature indications from the three thermocouples and the two Allen-Bradley resistors mounted in the torque tube. These graphs clearly show the dependency of the cold end heat leak on mass flow rate. Also, as the mass flow rate increases, the temperature profiles at 8000 rpm begin to resemble the curve at 1000 rpm. Hence it is possible to conclude that the torque tube performance is not rotationally dependent and that the torque tube temperature profile is only a function of mass flow rate. The cold end heat leak for the torque tubewas calculated from the temperature profiles of Fig. 4 and the associated mass flow rate . The one-dimensional coupled convection-conduction equation of the form (1) was solved for the temperature distribution between the known temperatures of the two Allen-Bradley resistors shown in Fig. 3. The conductance of the torque tube, including the stainless steel and Micarta, is essentially linear in T from 10 to 50 K, which allows for a closed-form solution of (1). Table I shows the calculated heat leak based on the cold end temperatures for each curve of Fig. 4. Also shown is the heat leak associated with the measured flow rate in the torque tube. The rhhrs product is corrected for the varying liquid Ievel in the rotor which in the presence of a centrifugal field creates a nonuniform distribution of thermodynamic states in the liquid system When the gradient heat leak and mass ftow rate heat leaks are equal, the torque tube flow is self-sufficient. This occurs at a mass ftow rate between 20 and 23 mg/s with an associated heat leak of 0.4 to 0.5 W. The magnitude of this heat leak is approximately as predicted by Tepper [ 1], and the associated temperature profile is nearly linear over most of the torque tube as predicted by Bejan eJ. eJ. Experimental Simnladon of a Cryogenie System for a Supereondueting Rotor 179 Table I. Beat Leak vs. Torque Tube Mass-F1ow Rate qL, W riJ,mg/s T1 T2 0.0 8.0 15.0 20.5 22.5 27.5 16.5 16.6 15.3 13.0 8.2 8.0 57.0 49.8 44.9 38.0 17.0 14.0 37.0 43.0 9.9 9.3 10.5 10.1 (1.1 = 8000rpm 10.0 6.5 4.6 2.8 0.27 0.11 w = 1000rpm 0.005 0.005 ±% mh1J0.95*, w 18% 18% 19% 20% 28% 35% 0.0 0.2 0.37 0.51 0.56 0.67 55% 58% 0.87 1.01 * Correction for work output of device. TableI also shows the value of vapor self-cooling. Self-sufficient flow reduces the cold end heat leak by at least a factor of 22 over the zero-flow torque tube heat leak. (In the zero-flow run the torque tube was colder than for the thermal steady state so the calculated heat leak is too low for this run.) To facilitate the self-pumping capability of the rotor eJ, the torque tube flow resistance should be small so that the pressure drop due to flow is small compared to the self-pumping pressure rise. The torque tube pressure drop has been recorded at 1000 rpm for mass flow rates ranging from 30 to 110 mg/s. The experimental relationship has the form In !l.P = 2.05 In m- 12.08 (2) which indicates a pressure drop of only 0.004 atm at self-sufficient flow. The pressure drop is a factor of 2 less than predicted C]. The slope of about 2 indicates that the pressure drop is associated with entrance and exit Iosses in the channels of the torque tube and in the rotating face seal [3 ]. At 8000 rpm the eflect of vapor compression prevents the accurate measurement of torque tube pressure drop. However, experience indicates that similar pressure drops are encountered at thesehigh speeds as weil. DRYOUT TESTS Experiments have been run to study the boil-ofl rate of liquid helium in the rotating frame. The experiments are necessary to provide a means of relating the mass flow rate of exiting vapor from the rotor to the heat leak into the reservoir. Estimation of heat leak based on the loss of latent heat in the fluid reservoir alone will be in error by as much as 25% at high tip speeds because the rotating reservoir is, in fact, an engine, converting some of the heat input into work as liquid vaporizes. Appropriate heat leak calculations during boil-ofl must take into account the loss of angular momentum within the rotor, the geometry of the liquid-filled spaces, as weil as the changing internal and kinetic energies of the helium in the rotor. A practical analysis of boil-ofl in a rotating frame requires making certain assumptions about the state of the fluid within the rotor. As liquid boils away, the vapor-liquid interface moves radially outward. The saturated vapor at the interface condenses as it migrates from a high pressure at the interface to a low, presumably fixed pressure at the centerline of the rotor. The radial pressure gradient in this - L Sollel, J. L s.ldl, Jr., _.. F. a-n "rainstonn" is established by the centrifugal compression of the vapor. The liquid droplets are thrown up against the radial baftles in the reservoir (not shown in Fig. 1) and form a slow-moving thin film. At each radins the centrifugal forces on the liquid are assumed to be balanced by the radial pressure gradient established by the vapor and the viscons shear between the liquid film and the baffte walls. The goveming equation of state in the two-phase region is therefore p., = J Jdp liJ 2 (3) rdr A single-phase region extends from the vapor-liquid interface to the outer radius of the winding space. In the absence of any heat leak, the fluid is compressed isentropically under the inftuence of centrifugal acceleration. A radial temperature gradient is established that is in convectively neutral equilibrium. However, heat conducted into the reservoir at the outer radins produces an additional radial temperature gradient which gives rise to gross convection currents transporting warm and less dense helium to the vapor-liquid interface where it is cooled by vaporization. The bulk fluid motion is assumed to be adiabatic though highly irreversible. With this assumption, the goveming equation of state in the singlephase region is (after Thullen [8 ]) r2 J'2 1 ,, j, dht- 2 = liJ 2 ds1-2 = 0 rdr, (4) With the previous assumptions, the first law equation written for a control volume encompassing the fluid reservoir and winding space and cutting through the torque tube of the rotor is . Q a +-J. at c.v. (liJ r)pdV = mh 2 2 0 a +at (1),~ (u +2 - pdV c.v. J. 2 (5) where m = -(ajat) Jc.v. pdVfromcontinuity. Thesecond termon theleft-handside of (5) accounts for the work produced by the control volume as a result of the torque from the loss of angular momentum due to boiling. This equation is integrated numerically using the NBS Helium Properlies Program. In order to evaluate the integral over the control volume, the states at incremental radü within the rotor are determined using (3) and (4). The rotor heat leak rate is assumed to be independent of time. Dryout tests were run, simply, by filling the rotor with liquid at 8000 rpm, shutting oft the inlet ftow, and then monitorlog the ensuing boil-o:ff history as the liquid Ievel in the rotor dropped. Each experimental run bad a different torque tube outlet flow rate which was held constant for the duration of a run. This insured a constant torque tube heat leak. The bulk of the exiting vapor left through the emergency discharge line. Figure 5 shows the mass flow time history of one such run. Superimposed on the experimental curve is the numerical solution of (5) for a (} roughly equal to that measured at the cold end of the torque tube. Tbe results of several tests were equally favorable and are described fully in other references [5 ' 7 ]. Tbe e:ffect of geometry on the thermodynamics of boil-o:ff in the rotating frame can also be seen in Fig. 5. Tbe step change drop and rise in mass ftow rate occurs as the vapor-liquid interface moves through the narrow radial channels which join the reservoir to the winding space. Experimental Simulation of a Cryogenic System for a Superoondncting Rotor 0 0 N .. "'- THE CIR ETICA L • -EXPERIME NT (fl~ -....- 281 • <..:> ::i:O "' w." 1-- N er.- a: 0 :J::O o....J LL..IIl ,... (/) (/) cr_O :r::"' "'"' 0 0 6 9 12 15 18 TI ME, MIN 21 24 27 30 Fig. 5. Theoretical and experimental mass ftow rate histories for constant heat leak at rotor outer radius. FLOW INSTABILIDES A ftow instability occurs at 7000 to 8000 rpm when the transfer tube inlet valve is shut and the torque tube and emergency discharge lines are unthrottled. The torque tube mass ftow will cycle from 0 to 40 mg/s over a period of 30 s. While the torque tube ftow is increasing, the emergency discharge ftow is zero. It was further noted that the centerline temperature cycles nearly in phase with the mass ftow out of the torque tube. Shortly after the centerline temperature reaches a maximum, the torque tube ftow drops to zero and gas begins to exit from the emergency discharge line. A probable explanation for this instability is as follows. When the torque tube outlet pressure is atmospheric, the isentropic compression of cold vapor in the torque tube inlet creates a vacuum at the centerline of the rotor. This centerline vacuum prohibits the ftow of vapor out of the emergency discharge or bayonet bleed lines. However, if the vapor at the torque tube inlet is heated, for example, by heat leaking down the external connections to the centerline of the rotor, there will be a progressive decrease in vapor density and a concomitant decrease in torque tube inlet vapor compression. Since the torque tube has a finite pressure drop, this slowly Ieads to a loss of centerline vacuum and of torque tube mass ftow. The loss of torque tube mass ftow results in an increased heat leak through the torque tube and into the fluid reservoir. The ensuing boil-off pressurizes the rotor until the centerline is slightly above atmospheric pressure. Then the boil-off gas can escape the rotor through the bayonet bleed and emergency discharge lines. This cools the centerline and the torque tube inlet in the process. Within seconds the density of the vapor column in the torque tube inlet has increased sufficiently to restore the centerline vacuum and the ftow through the emergency discharge line stops. This instability has significant implications. In the "self-pumping mode," that is, when one is relying on cold vapor compression in the torque tube inlet to lower the centerline pressure and draw in liquid from the dewar, the presence of a centerline heat leak may cause a loss of centerline vacuum. In essence, centerline heating could 282 L. Sobel, J. L. Smith, Jr., and F. Rumore cause a cessation of ftow at a time when greater ftow in needed. Similarly, a surge in boil-off could draw in too much liquid, causing liquid to spill into the torque tube inlets causing overpressurization in the torque tube and structural damage. A very significant source of heat leak into the reservoir comes from the transfer tube itself. An experiment was performed to est~tblish the magnitude of the heat absorbed by liquid helium as it travels from the dewar through the transfer tube and into the centerline of the rotor. The transfer tube was put into a partially filled liquid helium dewar so that the inlet of the tube remained suspended in the vapor phase above the liquid. The outlet of the transfer tube was connected to a liquid nitrogen shielded dewar which was connected at the outlet port to a mass ftow meter. The temperatures at the inlet and outlet of the transfer tube were measured by means of Allen-Bradley resistors mounted inside. The pressures at the transfer tube inlet and outlet were also measured. The heat leak in the transfer tube is calculated from the mass ftow and the measured temperatures and pressures at the inlet and exit. (The NBS Helium Properties Program was used to calculate the inlet and exit enthalpies.) The results are shown plotted in Fig. 6. The wide discrepancy in heat leak at the low mass ftows results from the inaccuracy in mass ftow measurement. The graph shows that the magnitude of the transfer tube heat leak is about 1.5 W. An error analysis establishes an uncertainty Iimit of ±6%, or ±0.1 W. For purposes of comparison this heat leak is nearly a factor of 3 higher than the self-sufficient coolant requirement of the torque tube. Thus a total ftow of approximately 70 mg/s is required to transfer any liquid into the rotor at all. The calculated heat leak is also much higher than the predicted heat leak for the transfer tube. However, some fraction of the calculated heat leak is attributable to end effects which arenot present when the transfer tube outlet is inserted into the rotor. To impede the convection of warm gas along the centerline from the warm parts of the transfer coupling to the helium reservoir, a set of baffies was inserted in the centerline helium inlet tube. The baffies consist ofthin paper disks cut to match the diameter of the centerline tube. The disks were skewered on a thin stainless steel welding rod and separated by spacers made of plastic tubing. The baffie assembly is removable. The edges of the paper disks are snipped around the circumference to allow a space for liquid to enter during fill-up and for cold gas to exit as weH. The additional pressure drop associated with these baffies is insignificant when the fluid is (.(')~,-------------------,~ 34.7 moJs •l. Iiter/ti' of l He , 4.2 K <!>~ a: 3: "' a:.n W-' ..J .... a: w :r: 10 20 70 60 so ~0 MRSS FLClH • MG/S 30 80 90 Fig. 6. Results of transfer tube heat Ieak test. 100 Experimental Simulation of a Cryogenic System for a Superconducting Rotor 283 very cold. The bafftes act as a one way valve. They facilitate the flow of cold gas in or out at the periphery of the centerline tube but blockwarm gas from coming down the center. To further insure that the centerline heat leak did not seriously affect the "dryout tests," a mass flow regulator was installed on the torque tube outlet. In essence, flow stability was achieved by placing a large flow resistance on the torque tube outlet. The dryout tests, however, are run with the inlet valve shut. Because of the centerline heat leak, it may not even be advisable to run the system in the steady state with a significant centerline vacuum when the inlet valve is open. CONCLUSION "Dryout" testing has established a frame of reference for studying the steadystate behavior of the cooling system. Future tests will determine two heat transfer coefficients. The first relates surface heater power to heater surface temperature and local winding space fluid temperature. The second is a gross convection heat transfer coefficient. This coefficient will relate heater power to heater temperature and reservoirfluid temperature minus the a T of isentropic compression. Furthermore, a model has been proposed describing complex cellular flow patterns convecting heat from the outer radius to the liquid-vapor interface. Verification of the theory depends on experimentally determining the thermal time constants of the fluid system. Finally, a major scope of testing will determine how best to control the effects of the centerline heat leak. The degree of torque tube and transfer tube throttling required to insure stability in the steady state will also be determined. NOTATION A = Jc.v.= cP = h = h18 = h0 = k(T) = m= P = 6= qL = ·, = s= T = T1 = T2 = u = W= x= conduction area control volume intergral helium specific heat enthalpy latent heat centerline enthalpy thermal conductivity helium mass flow rate pressure heat input cold end heat leak radius enthropy temperature temperature indicated by coldest Allen-Bradley in torque tube temperature indicated by next coldest Allen-Bradley in torque tube internal energy work output conduction length Greek Symbols w = rotational speed p = fluid density Pv = saturated vapor density o/ot =time derivative L. Sobel, J. L. Slllith, Jr., IUitl F. Rumore REFERENCES 1. K. A. Tepper, S.M. thesis, Massachusetts Institute of Technology, Cambridge, Massachusetts (1976). 2. A. Bejan, Ph.D. dissertation, Massachusetts. Institute of Technology, Cambridge, Massachusetts (1974). 3. K. A. Tepper, J. L. Smith, Jr., and F. C. Rumore, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 118. 4. D. Otten, "Demonstration of an Advanced Superconducting Generator," ERDA Contract No. E(49-18)-2295, Task Order No. 11-IR4, Interim Report (January 1978), p. 123. 5. L. Sobel, S.M. thesis (unpublished), Massachusetts Institute ofTechnology, Cambridge, Massachusetts (1979). 6. R. A. Bukovich, S.M. thesis, Massachusetts Institute of Technology, Cambridge, Massachusetts (1978). 7. L. Sobel and J. L. Smith, Jr., "Analysis of Boiloft of Helium in a Centrifugal Field," paper submitted to Cryogenics. 8. P. Thullen, Ph.D. dissertation, Massachusetts Institute of Technology, Cambridge, Massachusetts (1969). F-3 ROTOR COOLING SYSTEM FOR A 10-MVA SUPERCONDUCTING GENERATOR* M. T. Brown, M. E. Crawford, and J. L. Smith, Jr. Massachusetts Institute of Techno/ogy Cambridge, Massachusetts INTRODUcriON The Cryogenics Laboratory and the Electric Power Systems Laboratory at MIT are currently involved in an effort funded by the U. S. Department of Energy to develop a 10-MVA superconducting generator. This machine incorporates the most advanced concepts available in order to demoostrate the greatest possible growth in the state of the art. A rotor and stator comprise the generator. Within the rotor are the superconducting field winding and two electromagnetic shields. The winding and shields are kept at liquid helium temperature by the rotor cooling system. The rotor is supported between thin-walled, vapor-cooled torque tubes for thermal isolation, and further isolation is obtained by using an evacuated stator bore containing a stationary thermal radiation shield. The rotor composition is depicted in Fig. 1. The superconducting field winding is contained between two torque-carrying support tubes, and the winding is immersed in liquid helium supplied from the field-winding reservoir. The first electromagnetic shield, a highly conductive copper can shield, is mounted on the outside of the outer support tube. Grooves in the tube form liquid helium cooling channels for the can shield. The second shield is a damper winding of finely stranded and transposed wire. lt is separated from the can shield by a sheath of low thermal diffusivity material, and the winding is cooled by three layers of liquid-helium-filled tubes. The shield channel and damper cooling tube network are connected to a reservoir and separator to form a cooling system that operates independently of the field-winding cooling process. In steady-state operation the magnetic field is constant in rotor coordinates, except for space harmonics and phase unbalance of the armature-winding reaction field. These ac fields are screened from the superconducting field winding by currents induced in the damper shield winding and by eddy currents induced in the conducting can shield. The resulting power dissipation in the shields is the principal thermal Ioad that must be removed by the rotor shield cooling system. The system also serves to remove the transient heat dissipation within the shields during a generator powergrid fault. In this paper an analysis and performance predictions are carried out for the 10-MVA rotor shield cooling system during steady-state operation. * Work supported by U. S. DOE under Contract No. EX-76-A-01-2295, Task Order No. 11. 285 286 M. T. Brown, M. E. Crawford, and J. L. Smith, Jr. WOUNO OAMPER --~ STATOR CORE CONOUCTIVE CAN SHIELO ARMATURE TOR BORE TUBE F IELO WI NO I NG THERMAL RAD I ATION SHIELO Fig. 1. 10-MVA rotor and stator assembly. COOLING SYSTEM PROCESS The proposed damper shield and can shield cooling loop are shown in Fig. 2. Helium is fed into the field-winding reservoir. From there, it spills into the damper and shield reservoir, then cycles through the wound damper, into the separator, and back through the can shield. The boil-off gas generated in the cooling process is used to cool the rotor torque tubes. In the design there is a simple check valve located Fig. 2. Rotor cooling system and helium flow circuit. Legend: 1. emergency discharge, 2. field-winding current Ieads, 3. helium inlet, 4. field-winding reservoir, 5. field winding, 6. reservoir fill tube, 7. check valve, 8. separator, 9. conductive can shield, 10. shield and damper reservoir, 11. gas vent, 12. torque tube, 13. wound damper. Rotor Cooting System for a 10-MVA Supereonducting Generator 287 between the cooling cycle and field-winding reservoir to prevent the latter from pressurizing during a generator fault. In the steady state, fluid ftows from the reservoir to the damper where it absorbs heat and expands. Thus, on its radial return to the separator, it has a lower density than in the damper and shield reservoir, and this difference in density provides a pressure differential for ftow through the damper tubes. The separator will accumulate liquid by this process until the liquid pressure head in the separator exceeds that of the reservoir, at the shield radius, inducing liquid ftow from the separator to the reservoir via the shield, thus completing the cooling loop. During a generator electrical fault the increased heat dissipation in the wound damper and can shield will cause a rapid expansion of the helium. This will effectively stop the steady-state ftow loop, and cause blowdown whereby the helium in the damper and shield blow out into the reservoir and separator. In this pressurized state the check valve will seal oft the damper and shield reservoir, and the continually expanding heliumwill begin venting through the emergency discharge. The reservoir and separator will be sized to handle the transient cooling requirements. Refilling will take place after recooling of the shield and damper. ANALYSIS The objectives of the analysis were to determine the mass ftow rate of helium in the loop, temperature rises across the damper and shield, and the boil-off vapor rates which provide coolant to the torque tubes. Parameters in the analysis included the helium centerline condition, geometry of the cooling loop, and the power dissipation in the damper and can shield. The loop for steady-state cooling consists of a reservoir, damper cooling tubes, separator, and can shield cooling channels. The loop is modeledas a closed circuit, ignoring the slight changes in mass ftow through the circuit due to boil-off in the reservoir and separator. In fact, the boil-off rate is about three orders of magnitude less than the loop ftow rate, and the slight differences in mass ftow rate through the loop are insignificant. The model for this analysis is shown in Fig. 3. m • Vs-;11 . - ~EPARATOR 5 .. ~ - "" __j_ La R1•4.76 CM ~SATURATED ., - APOR ~~V TO TOfiOUE QRESERVOIR "" • TU BEmyR lq ~ LIQUID SEPARATOR RESERVOIR I-CIR CUIT MAss FLO w . ~ . 3 '-1.2 ATM SAT. LI Q. 1--- ~AN,HIELD R6 • 17 CM -~ 6 7 "' Rz• 20 CM __1 2 OoAMPER Fig. 3. Steady-state heating model of the damper/shield cooling circuit. - M. T........ M. 1:. c..wror., _. J. L s.ldt, Jr. Determination of the helium mass ftow rate involves linking it to the pressure drops in the system. In the analysis it is assumed that frictional pressure drops occur between states 6 and 7 in the can shield and states 2 and 3 in the damper. All other frictionallosses are neglected. The pressure drops in the loop are modeled using a conventional head-loss formulation written in terms of the mass ftow rate: llP = L) ·2) ·2 L(I[J,. ( 2 ~ 2 + L K{ 2;A 2) (1) In the above equation, I is the friction factor for smooth-walled turbulent ftow and centrifugal and Coriolis eflects on I are ignored. For the Reynolds number regime applicable to the cooling loop, it is assumed that I= 0.184[:"r 0 2 · (2) The L/ D,. term represents channel ftow length divided by hydraulic diameter, or the L/ D equivalent for the elbows and headers. The K coefficient accounts for entrance and exit Iosses and abrupt expansion and contraction losses. The analysis assumes that ftow in all radially directed pipes is one dimensional and isentropic. In the generator, radial ftows would be assured by judicious use of bafftes and tubes. This is necessary to keep any tangential shear forces to a minimum. By conservation of angular momentum a free ftow of high mass ftows would be overly dissipative. The change in enthalpy for isentropic ftow is 2 llh = -w2( r2 0 2 - r-) I (3) where w is the rotorangular speed and r; and ro are the inner and outer radü of the ftow loop geometry. Heat rejection from the helium to the reservoir or separator is modeled as a condensing process whereby the incoming fluid is returned to its saturated liquid state. This is assumed to occur at the radius of the liquid Ievel in the reservoir or separator. The heat rejection is given by tlout = m(h2." - h ..,) (4) Note that the vapor mass ftow for each torque tube can be approximated by equating the heat rejection to the product of the boil-ofl vapor rate, m.,, and the heat of vaporization, h18• Thus the boil-ofl vapor rate becomes • • h2<1>- m.,=m h hsat (5) fg Heat input to the ftow loop is from electrical power dissipation in the damper or can shield. lt is modeled as a first law heat addition tlin = m(hout- hin) (6) Figure 4 shows a typical set of eight thermodynamic states for the cooling loop. The equations described in the preceding paragraphs were used to connect these states. State 1 is fixedas saturated liquid at radius r 1 • State 2 follows from (3) and a fixed r2 with s1 = s2. State 3 is determined by assuming mand using (1) and (6). State Rotor Cooling System for a 10-MVA Superconducting Generator • 289 AHoAMPER • 2 liH= F(AR) H SATURATED LIQUID LINE AHSEPARATOR t s Fig. 4. Enthalpy--entropy diagram of fluid ftow states. 4 follows by assuming r4 in (3) and s3 = S4. State 5 is saturated liquid at T 4 • State 6 follows from (3) and a fixed r6 with s5 = s6 • State 7 is determined using (1) and (6). Finally, state 8 follows from (3) and a fixed r7 with s7 = s8 • At this point the saturated liquid condition corresponding to T 8 is compared with state 1. lf there is no agreement, a new set of rit and r4 must be assumed, and the solution is iterated to convergence. COMPUTATIONAL METHOD A unique computational procedure was developed to determine the mass ftow rate for a given set of parameters. Equations connecting the eight thermodynami c states were written using (1) through (6). An additional set of algebraic equations for the thermodynamic properties of helium were obtained by curve-fitting the The resulting set of 14 nonlinear coupled equations is given in NBS helium data Table I. These equations were algebraically combined using MACSYMA [ 2 ], a computer system for symbolic manipulation. The resulting equation is of the form eJ. r! + /2 (rit)r~ + [I(rit) = 0 (7) M. T. Browa, M. E. Cnwford, aad J. L. Smitla, Jr. 290 Table I. System of Equations Goveming Cooling Loop in Steady-State Mode Defining Equation qd = m(h3 - hz) Pz- P 3 = d'm ~.s + e'm 2 = a'h 3 + b's 3 + c' h 3 - h 4 = w 2 /2(r~- r~) h4 = fT 2 + T(g' + SJ) + i' h 5 = j'T+ k' P3 s = l'T+ o' ~- h 5 = w 2 /2(r~- r:) State 3 Property State 4 Property State 5/Property q",. = m(h7 - h 6 ) State 6 Property State 7 h 1 - h 8 = w 2 /2(r~- ri) h 8 = aa's 1 + bb' Property State 8 State 8/Property P6 = u'h6 + v's 5 + w' p6- p1 = r'mL8 + t'm2 P1 = u'h 1 + v's 7 + w' Note: Defined system parameters h1, st> PI hz, Sz, Pz qcs, qd rb Tz, r6 where / 1 and / 2 are algebraic functions. This equation can be factored and rearranged to obtain (8) Examination of (8) shows that for real values, the minimum mass flow will occur when (9) Equation (9) was solved for musing a Newton-Raphson iteration method, and the corresponding r4 was obtained from (8). Additional values of m that Slltisfy (8) were determined by incrementally changing the mass ftow. The upper limit on mis the condition when either h 8 = h 1 or h4 = h 5 , corresponding to zero heat rejoction in either the reservoir or separator. Note that the quadratic nature of the tolution implies that there will be two values of r4 for a given m. RESULTS AND DISCUSSION Sampie calculations have been carried out for the geometry given in Fig. 3, based on the current 10-MVA design parameters eJ. State 1 was fixed at 1.2 atm saturated liquid at r 1 = 4.76 cm from the centerline. The heat inputs to the damper and can shield were qd = 9.2 W and 4cs = 1.6 W. The damper coolin& circuit consisted of 624 tubes, each with 0.284-cm2 ftow area, and the can shield cooling circuit consisted of 72 channels, each with 0.227-cm2 ßow area. Four sets of perforations through the support structures connected the separator/reservQj.r to the can/damper, and were a significant factor controlling the circuit mass ßow rate. In the sample calculation 12 ßow channels oomprised a set of perforations (i.e., 12 291 Rotor Cooling System for a 10-MVA Superc:ondncting Generator 1.180 1.178 1.176 1.174 P, atm 1.112 1.170 1.188 --RESERVOIR CENTERLINE _ _ __.,__ PRESSURE 1.168 138.1 1584 1585 138.2 1385 138.8 138.7 138.8 ril . g/s Fig. 5. Pressure at the centerline in the separator vs. circuit mass flow. channels connecting the reservoir to the damper). On the reservoir end the total perforation cross-sectional flow area was set at 50% of the total can and damper cross-sectional area. On the separatorend the perforation area was 100%. Figure 5 shows a graph of the centerline pressure of the separator vs. mass ftow rate for a given reservoir centerline pressure. Note that the minimum mass flow is 138.7 g/s and that there are two values of possible balancing pressures (i.e., two r4 values) in the separator for each mass ftow. Figure 6 shows the heat rejected from the reservoir and separator for each of these two values and the end points are defined by • 8 QREJECTION~ Watts / / 8 / ./' -- _........---'- - - - =- ""'---SEPARATOR ~~-( 4 3 2 \. " RESERVOIR .......... ............ .............. ......_ 138.6 138.8 ril, g/s Fig. 6. Heat rejection in the separator or reservoir vs. circuit mass flow. M. T. Bnnra, M. E. Cnwfonl, ad J. L Smltll. Jr. either zero heat rejection or total heat rejection in the separator (or reservoir). By using (5), the variation of vapor mass ftows into each torque tube can be calculated. lt should be noted that for actual operation of the generator it would require control of both the reservoir and the separator centerline pressures to achieve the desired Operating state. (This would be by either torque tube pressure controls or controls of the boil-ofl-vapor mass ftow rates.) Calculations with variations of the reservoir centerline pressure on the order of ±0.5 atm (at R 1 = 4.76 cm) produced no significant changes in the nature of the operating curve in terms of shape or mass ftow rates. Thus the reservoir centerline pressure may be made to vary with the understanding that only the magnitude of the separator pressure (referred to the reservoir) would vary on the depicted curves but they still would relate the generat variance of the second input, separator pressure (or vapor-mass ftow rate), for the system Operation. The sample calculations indicate that the natural-convection-driven helium mass ftow induced by the heat input is large, resulting in small temperature increases across the can and damper. For the range of mass ftow rates in Fig. 5, the average temperaturein the can is about 4.75 K andin the damper it is about 4.88 K, and the temperature rise is about 0.05 K. A calculation was also performed for the situation involving a doubling of the heat inputs. For this case, the damper temperature rises are about 0.10 K. SUMMARY A rotating closed-loop liquid helium cooling system has been designed that derives its pumping action from a natural convection or thermosyphon effect. The loop is comprised of two axially oriented warm reservoirs which receive heat from the electromagnetic shields, and annular cold reservoirs at either end of the warm reservoirs which dissipate the heat by boiling. The major driving mechanism is the centrifugally created pressure diflerence across the outermost warm reservoir, with a warmer radial column of helium at one end compared to the other end. The loop pumping capacity is quite high. For the sample calculation the self-pumping rate is about 139 g/s for a 10.8-W heat input to the warm reservoirs. By changing the annular reservoir geometry the helium mass ftow rate can be adjusted to accommodate frictional pressure drops associated with various proposed header designs that connect the warm and cold reservoirs. NOTATION A = cross-sectional area D,. = hydraulic diameter f = friction factor in head-loss equation h = enthalpy h,8 = heat of vaporization K = coefticient in head-loss equation L = length m = mass ftow rate P = pressure q = heat ftow rate r = radii from centerline s = entropy T = temperature a', b', c', etc. = all primed letters are coefficients for either curve fits or parametric relations Rotor Cooling System for a 10-MVA Superconduding Generator 293 Greek Symbols p = density w = rotor rotational speed REFERENCES 1. R. D. McCarty, NBS Tech. Note 631 (1972). 2. MACSYMA Reference Manual, Math. Lab. Group, Project MAC, Massachusetts Institute of Technology, Version 9, December 1977. 3. "Demonstration of an Advanced Superconducting Generator," Interim Rept. 7, Massachusetts Institute of Technology, Cryogenic Engineering Laboratory and Electric Power Systems Engineering Laboratory, E(49-18)-2295, Task Order No. 11, IR 7 (1979). F-4 SAFETY LEADS* M. Kuclmir and T. H. Nicol Fermi National Accelerator Labaratory Batavia, Illinois INTRODUCDON The current plan for the Fermilab superconducting synchrotron (energy doubler) calls for the installation of approximately 1000 superconducting magnets in the 6-km tunnel of the main ring accelerator.lts 3 x 108 -J emergency energy dump system CJ is based on "safety Ieads" between 4 and 300 K situated after every fifth magnet. With this system, when a quench is detected the power supply is turned off and a 0.5-!l air-cooled "energy fountain" resistor is switched into the coil buss at each of six energy transfer stations. This causes the magnet current to decay with a time constant of 10 s. This decaying current has to be diverted from the developing quench, and this is done by shunting the current out from the group of five magnets containing the quench by means of the "safety Ieads." Protection for these magnets is achieved by firing their internal heaters which spread the normal zone to a safe size. The magnetic energy: of these five magnets is transferred to the helium and mechanically vented eJ. DESIGN CONSIDERATIONS The above-described particular use for these current Ieads and the large number required justify an optimized design. Considerable capital savings in room temperature plumbing is achieved by dry (instead of vapor cooled [3 ]) Ieads. Advantage can be taken from the infrequent nature of their use to minimize their inactive state heat leak Ioad. The Iead itself can be allowed to become bot when carrying current, but its junction to the magnets must remain superconducting otherwise the quench will propagate out of the confined five-magnet cell. To satisfy this condition the coldest part of the Iead (the "quench stopper") has a large contact area to liquid helium. High valtage to ground (several kilovolts) might develop between the Iead and the cryostat, requiring good electrical isolation and making heat sinking to 78 K impractical. The Iead will conduct heat directly from 300 to 4 K through the thermal insulation vacuum space. Several materials were considered for the bot section of the Iead. Numerical calculations for the expected temperature rise of a small segment of Iead as a function of time were carried out for these materials under an adiabatic approximation. In this approximation, the rate of temperature increase, dT/ dt, of an element of cross section s and unit length of the Iead is determined only by the * Work sponsored by the U. S. Department of Energy. 194 295 Safety Leads electric power, / 2 p/ s, dissipated in the lead and its heat capacity IJ.CS (1) where p(T), c(T), and IL are the resistivity, specific heat, and density of the material, respectively. In this approximation cooling effects are neglected. The maximum expected current of 4.6e -t/lD kA (where t is in seconds) will cause the element to reach a maximum final temperature, T max• which depends on its initial temperature, Tinit• and the quench load as given by (2) A new material property, the quench load capability f, can be defined as - F- f- 2 - 1L S f Tmax Tm&t c(T) P (3) (T) dT For maximum quench load per heat leak, one should maximize the ratio F 6 s 2f s/ I J~ 00 k(T) dT = (4) slz where k(T) is the thermal conductivity and z = f!U:oo k(T) dT] is an index-of-merit of the material. Table I presents calculated values for fand z of selected materials for a conservative 1init of 300 K and a practical T max of 514 K. Constantan was selected instead of stainless for its electrical characteristics as weH as for its availability as a well-characterized material. The value of f in equations (2) and (3) specifies a value for s of 2.5 cm 2 • This cross section, the specification of an allowable heat leak of 0.5 W and the thermal conductivity integral of Constantan (516 W /cm from 4 to 300 K) determine a value for I of 258 cm. The known resistivity of Constantan (3.6 x 10-5 n cm) permits an easy estimate of the maximum voltage drop (4.6 x 103 x 258 x 3.6 x 10-5 /2.5 = 17 V) across the lead, a value needed for electrical considerations. Table I. Load Capability and Index-of-Merit of Selected Materials for a Heat Excursion from 300 to 514K f, Material MJ/Ocm Copper Nickel Constantan Stainless Niobium Titanium 314 56 17.2 10.4 22 8.4 4 Z, s/~-tficm 3 0.20 0.26 or 0.09 0.33 0.34 0.14 0.08 M. "--lllir Md T. H. Nicol QUENCH STOPPER DESIGN CONSIDERADONS The liquid-helium-immersed part of the Iead cannot be treated in the same adiabatic approximation; rather, a term expressing the heat transferred to the liquid is essential. This is provided by dT p(T) 2 p,sc(T)-d = I ( t ) - - hp(T- TL) t s (5) where h is the heat transfer coefficient, p is the perimeter of the quench stopper cross section, and TL the temperature of the liquid. Numerical solutions of this differential equation for several materials using heat exchange coefficients found in the Iiterature [4 ] indicated that a copper conductor/radiator 25.4 cm long with an 8-cm2 cross section and a surface area of 5000 cm 2 was suitable. These calculations were based on copper properlies that will probably be different for the copper actually used. To overcome this difficulty an actual test of the performance of such a quench stopper was carried out and is described after the construction details. CONSTRUCDON DETAILS A five-magnet unit is composed of four dipole magnets and one quadrupole magnet. The quadrupole magnet cryostat contains a service volume to house various components necessary for the operation and protection of the system. Relief lines for the helium and nitrogen shield, instrumentation, correction coil power Ieads, vacuum relief and pumpouts, beam sensors, and transition piping are among those services contained therein. The available space for these components is approximately 36 cm high, 46 cm wide, and 51 cm long. The safety Iead is also incorporated in this area. Its drypartwas formed from a cable of Constantan wire to a shape compatible with the available space. Cable was chosen over solid stock for ease of forming and availability. Nineteen strands of 0.411-cm-diameter wire were needed to satisfy the required cross section. This cable was then formed to the shape shown in Fig. 1. Electrical insulation is provided by one layer of Kapton followed by one layer of glass cloth tape. Connection at both ends, one to the quench stopper, the other to an extended cable, is via ceramic vacuum feedthroughs. Internat support and standoff is by G-lO saddle pieces. As previously indicated the quench stopper is a copper conductor with large heat exchange area. To comply with the low profile required by space limitations, this device was made of readily available copper strips 25.4 cm long, 1.27 cm high, and 0.079 cm thick. These were assembled as shown in the insert of Fig. 2, stacked 80 across and fumace brazed to form the assembly shown in Fig. 2. The resulting unit is 25.4 cm long, 1.27 cm high, and 12.7 cm wide. Unlike the safety Iead .cable which lies in the insulating vacuum space, the quench stopper is in a space ftooded with single-phase liquid helium (see Fig. 1). During a quench, potentials as high as 5 kV could develop between this assembly and ground (the single-phase box). lnsulation must be provided which will not only withstand this potential difference but will allow the free ftow of helium across the "fins" of the quench stopper. Safety Leads 297 .162" Dia Constontan 19 Ploce5 SINGLE PHASE HELIUM SPACE OUENCH STOPPER LEAD CABLE Fig. 1. Safety Iead shape and location in the service volume of the quadrupole cryostat. NEMA G-10 pieces surround the finished assembly at all points where it is in close proximity to the walls of the single-phase box, as shown in Fig. 1. The placement of the quench stopper is in an active cooling region, i.e., directly over the relief line opening. Relief valves are activated at the time of quench detection to prevent overpressure. Helium ftowing past the quench stopper fins maintains the splice end of the assembly below the superconducting transition temperature. -• - CORNER DETAIL SHOWING ALTERNATING 1/32" STRIP 1/32" SPACE I Fig. 2. Quench stopper section of the safety Iead. M. Kuchnir and T. H. Nicol 298 SUPERCON CABLE Fig. 3. Quench stopper test setup. QUENCH STOPPER TEST The first prototype was instrumented with precalibrated carbon resistor thermometers: T 2 near the Constantan junction and T3 near the superconducting cable splice. This prototypewas installed after a vacuum-encapsulated 76-cm-long section of Constantan cable. Care was taken to get a reasonably good simulation at the junction between the Constantan and the quench stopper. To insure that T 2 and T3 measured the temperature of the copper, the thermometers and some immediate length of their CuNi alloy Ieads were varnished in a hole drilled in the copper and covered with epoxy to isolate them from the liquid. Another carbon thermometer, 18 17 FINAL T1 > 460 K l>T3 <50 mK T l~rouq~ou l ~~~I (Q.,I ~ 2 !3 U I (Q (Q3) I 0 10 20 30 40 50 TIME , sec I I I 60 70 80 l Fig. 4. Temperature excursions for three current pulses in the quench stopper test. Safety Leads 299 Tt. monitored the temperature of a specific location in the Constantan. A superconducting cable completed the circuit. Figure 3 shows this arrangement. Seven pulses of current simulating quenches, 01 through 07, were fired. Figure 4 shows events 03, 06, and 07 with the resultant temperature excursions of T 2 (top of quench stopper) and T3 (bottom of quench stopper). 06 corresponds to a typical maximum quench and 03 to common weaker quenches. 07 was a finallong-held pulse, which also bad minimal effect in raising the temperature of the bottom of the quench stopper. ACKNOWLEDGMENTS The authors would like to acknowledge the supportof G. Kalbfteisch, G. Biallas, N. H. Engler, and H. Fulton in the development, and W. A. Wojak, H. Warren, W. Habrylewicz, C. Hess, and J. Tague in the testing. NOTATION c = specific heat F = quench Ioad [see equation (2)] f = quench capability [see equation (3)] h = heat transfer coefficient I = electrical current k = thermal conductivity I = length of conductor p = perimeter of cross section () = heat leak Q1-Q7 = events simulating quenches s = cross section of conductor t =time T = temperature 1init = initial temperature of a Iead section TL = temperature of liquid helium T max = maximum temperature of Iead T 1 - T 3 = thermometers used in the test or their temperatures z = index of merit Greek Symbols p. p = density = resistivity REFERENCES 1. 2. 3. 4. R. Stiening, R. Flora, R. Lauckner, and G. Tool, IEEE Trans. Magn. Mag-15:670 (1979). M. Kuchnir and K. Koepke, IEEE Trans. Nucl. Sei. NS-26:4045 (1979). M. C. Jones, V. M. Yeroshenko, A. Starostin, and L. A. Yaskin, Cryogenics 18(6):337 (1978). P. J. Giarratano and R. V. Smith, in Advances in Cryogenic Engineering, Vo/. 11, Plenum Press, New York (1966), p. 492. DISCUSSION Question by M. A. Green, Lawrence Berkeley Laboratory: Did you consider 304 stainless steel as a Iead material? How did it compare with the Constantan material? Answer by author: Yes, we did.ltwouldresultin a thicker andshorter Iead, quite acceptable. Butforthis prototype, we decided on Constantan for its smaller temperature coefficient that obviated some transient analysis on the voltage drop calculation. F-5 MAGNET LEADS FOR THE FIRST-CELL* D. P. Brown and W. J. Sdmeider Brookhaven National Laboratory Upton, New York INTRODUCDON The ISABELLE refrigeration system utilizes compressed liquid helium to supply refrigeration to nearly 1100 superconducting bending and focusing magnets. These magnets steer the proton orbits of the accelerator and are arranged into two interlocking rings. The magnet Ieads used to power, correct, and trim these superconducting magnets make up a substantial portion (25%) of the total heat Ioad that the refrigerator must be capable of supplying. These magnet Ieads have been designed to minimize the heat input into the refrigeration system. The design and preliminary test results of four different types of Ieads in use on the ISABELLE prototype, the first-cell, are described. DISCUSSION c- Magnet Ieads are designed to minimize 3 ] the heat input into the refrigeration system. This is accomplished by optimizing the cross-sectional area of the conductor and intercepting with some gaseous helium ßow, a portion of the resistive heat generated, and the heat conducted from the room temperature end into the coolant. The total refrigeration [4 ] cooling Ioad, Q" is the sum of the heat ßux, Q", from the cold end of the Iead into the coolant and the mass ßow, m, up the Iead times some factor, f, characterizing the refrigerator cycle. Thus Q, = Q" + mf (1) where f is a factor which characterizes the refrigerator in a refrigerator/liquefier mixed-duty operation. For a theoretical Carnot cycle plantat 4.4 K this factor would have a value of 80 W-s/ g and, for actual refrigerators, values in the range of 50 to 95 W-s/ g have been reported [4 ]. The actual value depends on the relative efficiency of the plant when operated as a refrigerator vs. operation as a liquefier. For the particular refrigeration cycle that will be used for ISABELLE, in mixed-duty operation approximately 630 W of room temperature input power is required to produce and deliver 1 W of magnet refrigeration below 3.8 K and 33,700 W are required to * Work performed under the auspic:es of the U. S. Department of Energy. 300 301 Magnet Leads for the First-Cell ELECTRICAL FEEDTHROUGH FOR VOLTAGE TAPa TEMPERATURE SENSORS GAS FEEDTHROUGH ELECTR I CAL FEEDTHROUGH ME TAL SEAL F L ANGE - I --1 TEMPER AT URE SENSORS OFHC COPPER CONDUCTOR MAGNET LEAD VESSE L ~ ' ~ SEE FIG. 5 1_ - " ' t -:nu_ STA I N LESS STEE L SUPPORT TUBE SUPERCONDUCTOR BUSS Fig. 1. 4250-A main power Ieads. produce 1 g/ s of ftow for lead cooling. The value of f used to optimize the leads is the ratio of these two values, 53.4 W-s/g. CONSTRUCI10N DETAILS There are four different classes of leads for varying current ratings designed for the first-cell and ISABELLE. These are delineated in Table I. The current rating and 301 D. P. Brown and W. J. Schneider Table I. Classes of Leads for Varying Current Ratings Designed for tbe First-CeU and ISABELLE Oass Lead type Current rating, A II I Mainpower Sextupole quad trim Rohinsbypass ±4250 ±300 III Octopole decupole duodecupole ±50 X 12 IV Steering Steering ±50 ± 100 X X 12* 12 Number required for first-cell Calculated conduction heat Ioad, Calculated total heat Ioad, 2 4 4 1 4 4 4 4 0 6.1 0.7 19.1 1.8 0.8 2.6 0.7 1.6 1.7 5.2 w w *Not fully used. reason for the various Ieads required for ISABELLE have been discussed previously [s]. The 4250-A (Figs. 1 and 2) and 300-A Ieads are similar in design except that the main power Iead is made of OFHC (oxygen-free high-conductivity) copper, while the 300-A Iead is manufactured of brass. A helix (acme thread) is cut into the conductor (brass or OFHC) outer diameter. The conductor is then shrunk into a stainless steel support tube. Helium gas flows up the helix, and, as a result of the design, a steep temperature gradient exists in the conductor at the warm end from 50 to 300 K. At the outlet, the gaseous helium flows through an insulator which isolates the Iead from ground and returns the gas to the helium refrigerator compressor suction. The 50-A (Figs. 3 and 4) and 100-A Iead designs are similar to each other except for the size of the twelve conductors. A Teflon core has a helix (acme thread) machined into its outer diameter. Twelve OFHC copper wires are equally space<i Fig. 2. Actual 4250-A main power Iead. Magnet Leads for the First-Cell 12CONOUCTOR 50A LEAD ELECTRICAL FEEOTHROUGH GAS OUTLET-300 K ELECTRICAL LEA OT HROUGH FOR TEMPERATURE SENSORS METAL SEAL FL ANGE TEFLON SHEATH 12 INDIVIDUAL OFHC COPPER GONDUCTORS TEMPERATURE SENSORS TEFLON CORE Wl TH GAS HELIX HELIX GAS PASSAGE . __.-MAGNET LEAD VESSEL SEE FIG. 5 !"'-·----~ } __ __J GAS INLET 4 K Fig. 3. Twelve-conductor 50-A correction or steering Iead. 303 304 D. P. Brown aad W. J. Seimelder Fig. 4. Actual twelve-conductor 50-A correction or steering Iead. TEMPERATURE SENSORS TOTAL OF 6 EA . MAGNET LEADS PER VESSEL I I . ' I 3"0D MAIN COOLANT TUBE INDIUM SEALS Fig. 5. Magnet Iead vessel. 16" (40.6 cm) LONG x I" (2.54 cm) 0 0 TO HOLD LEADS TUBES Magnet Leads for the First-CeU 305 around the periphery of the Teflon core intercepting the helix. A Teflon jacket is then shrunk around the diameter of the Teflon core holding the wires in place. Heliumgas traveling around the helix intercepts heat in each of the twelve wires. Owing to this configuration, a steep 50- to 300-K temperature gradient results at the warm end in the copper wires. The gaseous helium then flows through an insulator and returns to the helium refrigerator compressor suction. Because of the relatively modest current rating (50 or 100 A), these single Ieads are capable of handling twelve separate circuits. Consequently, the 50-A lead has a total capability of 600 A and the 100-A Iead is capable of 1200 A.lt is not desirable to parallel any of these circuits because the current flow is not self-regulating. However, if a parallel scheme is used, it is essential that a control system be employed to insure that the current flow divides evenly. FIRST-CELL The nurober of Ieads as shown in Table I are required to meet the power requirement of the first-cell. The magnet lead vessel shown in Figs. 5 and 6 is designed to accommodate six Ieads. Therefore, two magnet Iead vessels are required for the first-cell. By way of comparison, ISABELLE will require upwards of 100 magnet Iead vessels. The heat leak to the 4-K stream on these vessels is expected to be5W. Fig. 6. Actual magnet Iead vessel under construction. 306 D. P. Brown and W. J. Schneider The two-magnet lead vessels provide the transition between the 4-K environment of the magnets and ambient temperature. Each of the first-cell leads is equipped with a temperature sensor which modulates an electric control valve or solenoid valve and maintains some preset temperature at one point along the length of the lead. The design of the control systein is described elsewhere [6 ]. Voltage taps are provided at the top and bottom of the leads. The differential voltage measured is a function of the integrated resistance of the lead and, for a given current, is an indicator of the temperature protUe of the lead. The differential voltage is incorporated into a safety circuit which will shut down the power supply in the event a preset voltage drop is exceeded. The steering dipoles (QF and QD) are powered from the tunnel end in a vessel that is integral with the magnet interconnector. One class 111 50-A lead is employed. However, it is not fully used. HEAT TRANSFER Numerical calculations of lead performance have been made with a computer program developed at CERN [ 4 ] and modified for the conditions at ISABELLE. These results were used to optimize the various class leads. Figure 7 shows a plot of the mass ftow and differential voltage vs. the current level for minimum refrigeration requirements on the 4250-A lead. z' 0 . 30 ~ 1<l D:: 0 . 25 loJ ~ > D:: LI.. loJ 0 . 20 10 0 D:: 0 D:: 0 ~ ::> - ::l; l - u 0 . 15 ~ lo.l 80 LI.. ...J 0 .1 60 > loJ 0 .05 40 ~ 0 ...J LI.. 0 0 <l loJ 1<l D:: loJ \!) <l 1- ~V) 0::::=; ol!> E Q.. 0 20 1000 LEAD 2000 3000 4000 50 0 0 CURRENT, A M PE RES Fig. 7. 4250-A Iead performance at minimum refrigeration. ...J Magnet Leads for the Fmt-Cell 307 The CERN design used on the main power and the trim Ieads provides an enhanced heat transfer area and a stabilizing thermal mass which inhibits aceidentat Iead burnout. The multibundle-type Iead used for corrections and steering is a new concept in Iead design. The multiplicity of these Ieads necessitates bundling to reduce the complexity of the control system and reduce space requirements. The heat transfer area for the 50- and 100-A Ieads appears tobe adequate. FABRICATION AND TESTING Leads from each dass have been fabricated and are presently being assembled into a magnet Iead vessel. ISABELLEdipolemagnet 005 is also being assembled for a force ftow test scheduled for late summer of 1979. Preliminary tests of Ieads similar The referenced 7-ft bubble chamber Ieads were to these have been excellent designed for and operated with hydrogen as a coolant. In addition, prototype CERN-type Ieads have been built and tested on earlier ISABELLE magnets using both helium pool boiling and forced-ftow cooling. Results for these Ieads have been consistently good and compared favorably to the predicted values although they operated somewhat cooler than anticipated. eJ. ACKNOWLEDGMENTS The authors wish to acknowledge the assistance of many members of the ISABELLEstaffin designing, fabricating, and testing these Ieads, and A. Walsh for graciously typing the paper. REFERENCES 1. R. K. Thomas, J. R. Purcell, and R. W. Boom, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 219. 2. C. L. Goodzeit, "Optimization of Gas Cooled Superconducting Magnet Leads-A Method for General Materials," BNL Report, BNL # 17233, (September 1972). 3. W. J. Schneider, Brookhaven National Laboratory, private communication. 4. D. Güsewell and E. V. Haebel, in Proc. 3rd Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1971), p. 187. 5. G. Parzen, IEEE Trans. Nucl. Sei. N26(3):3995 (1979). 6. M. Afrashteh, Brookhaven National Laboratory, private communication. DISCUSSION Question by M. A. Green, Lawrence Berkeley Laboratory: In the final figure of the paper, are the values of valtage drop and mass ftow given for a single Iead or are they given for a Iead pair? Are the Ieads designed specifically for your refrigerator, which appears to have a much lower value of refrigeration to liquefaction coefficient than most refrigerators? Answer by author: In the final figure of the paper, the valuesof valtage drop and mass ftow are for a single Iead. Yes, the mass ftow that the Ieads have been optimized for are specifically for the ISABELLE refrigerator. However, use of the Ieads on a different refrigerator only necessitates that the mass ftow be adjusted to produce minimum refrigeration. lt would be necessary to know [, the factor which characterizes the refrigerator/liquefier mixed-duty cycle to obtain that optimization. F--6 THERMAL CONTROL FOR THE MFTF MAGNET* J. H. VanSant and R. M. Ross Lawrence Livermore Laboratory, University of Califomia Livermore, Califomia INTRODUCDON The Mirror Fusion Test Facility (MFTF) will have the world's largest magnet. Externat dimensions of the assembled Yin-Yang magnetwill be 7.8 by 8.5 by 8.5 m, and it will weigh approximately 300 tons. More than 8000 Iiters of circulating liquid helium will be required to maintain the nearly 50 km of superconductor at below 5.0 K while the latter carries almost 6000 A in a magnetic field up to nearly 7.7 T. Such formidable conditions have necessitated extensive planning to achieve the required thermal control. This paper describes four features of these thermal control plans. First, it outlines the proposed cooldown and warmup schedules for the MFTF and the procedure for regenerating extemal cooling surfaces. Then it discusses the design of an extemal quench resistor, based on an estimate of the superconductor's maximum temperature. Last, it explains how a computer model of liquid helium circulation was used to aid in choosing pipe size for the liquid helium lines. To minimize heat Ioads on the magnet system, all liquid helium temperature surfaces will be in a 1.3-l.f.Pa (10-nTorr) vacuum and shielded by liquid nitrogen. In fact, the entire magnet, which will be maintained at less than 4.5 K, will be installed in a 12-m-diameter, 18-m-long vacuum vessel that also will be the mirror fusion chamber. The vessel will support the magnet by means of seven liquid-nitrogencooled banger and stabilizer rods. Externat surfaces of the magnetwill be shielded by 311 m 2 of liquid-nitrogen-cooled panels. In regions where plasma and neutral beam impingements will occur, 100m2 of water-cooled panels will shield the liqwd nitrogen panels. A sectional view of one magnet coil in its major radius region is shown in Fig. 1 and illustrates typical MFI'F construction. Tables I and II describe the magnet materials and conditions. A unique feature of the magnet design is the guard vacuum space, which serves two principal purposes. It protects the fusion chamber from developed helium leaks through the support structure and provides an additional ftow channel to circulate gas for more eftective cooldown or warmup of the structural case. Moreover, the vacuum space and the urethane shim help to insulate the magnet coil. • Work pcrformed under the auspices of the U. S. Department of Energy by Lawrence Uvermore Laboratory under Contract No. W-7405-Eng-48. 308 309 Thermal Control for tbe MFfF Magnet pla sma shield Neutra l beam shield Fig. 1. A section view of one MFTF magnet coil, showing construction details for both the magnet and for the cooling system that surrounds it. The section is along Iine AA in the small drawing. Table I. Properties of tbe MFfF Magnet Design current, A Maximum magnetic ftux density, T Conductor size, cm Copper-to-superconductor ratio Effective copper area for current, cm 2 Conductor surface area for heat transfer, cm 2 /cm Transition temperature at maximum current and field, K Maximum liquid helium temperature, K Number turns per coil Conductor length per coil, m Coil porosity Stored magnetic energy, MJ 5775 7.68 1.24 X 1.24 6.7/1 1.25 8.17 5.0 4.52 1392 24,785 0.32 440 310 J.H.V..s.tudR.M.R. . Table D. Materials aad Conesponding Weigbt for the MFI'F Mapet Materials Magnet Conductor (copper, superconductor) Insulation (G-11) lacket (316 stainless) Sbim (uretbane) Bladder (copper) Case (304 LKN stainless) Intercoil supports Gas baflle Total Accessories Liquid uitrogen shields (316 stainless) Water shields (316 stainless, copper) Liquidhelium ducts (316 stainless) Current Ieads (copper) Hanger rods (316 stainless) Miscellaneous hardware Total Weight, kg 52,820 5,730 16,180 3,370 1,130 178,370 36,360 6,180 300,140 10,650 3,430 2,730 500 12,730 1,100 31,140 COOLDOWN AND WARMUP The magnet cooldown schedule will be programmed to prevent unacceptable structural thermal stresses. Cooling gases circulated through the coil and guard vacuum spaces will be supplied at temperatures decreasing from 225 to 100 K during a minimum 50-hr period. During this period, approximately 20 g/s of helium and 400 g/s of nitrogenwill be circulated through the coil and guard vacuum spaces, respectively. Afterwards, the helium ftow will be increased to 100 g/s, and the guard vacuum will be pumped down to approximately 13 Pa. An additional 50 hr are estimated for complete cooldown to 4.5 K. The total energy extraction is approximately 29 GJ. Severe ftow asymmetry between the two coils (e.g., 30%) is not expected to cause unacceptable thermal stresses, but ftow control will be employed. During warmup of the magnet, the gas temperature schedules will be nearly reversed from cooldown. That is, the helium supply temperature will increase from approximately 75 to 150 K during a 50-hr period at a ftow of 100 g/s through only the conductor pack. Then nitrogen and helium will be circulated through the guard vacuum and coil spaces at supply temperatures increasing from 150 to 300 K during the next 70 hr. The design stress Iimit for the structural case is 550 MPa (80,000 psi) which provides a 1.5 factor of safety based on tensile yield strength. Calculated stresses during cooldown and warmup range from approximately 55 MPa (8000 psi) compression to 90 MPa (13,000 psi) tension resulting in a factor of safety greater than 9. A numerical analysis of the magnet cooldown and warmup cycles performed by General Dynamics Convair Division [ 1] yielded the foregoing thermal schedules. They are intended to prevent thermal stresses from exceeding design Iimits. However, strain and temperature at many locations will be. monitared during thermal cycles. Thermal Control for the MFI'F Magnet 311 EXTERNAL SURFACE REGENERATION Cryopumping panels in the MFfF vessel will periodically be thermally cycled to regenerate the surfaces. Tbe magnet's external surfaces must be warmed to approximately 15 K at the same time to prevent released gases from recondensing. To accomplish this, the magnet liquid nitrogen shields are to be drained and warm nitrogen gas (300 K) circulated through them to radiate heat to the magnet surface. The shield temperature will be controlled so as to minimize cumulative heat transfer to the magnet. Prior to the regeneration cycle, the magnet must be deenergized to prevent a quench. Subsequently, the liquid helium must be pumped out so that transfer lines inside the vacuum vessel can also be regenerated. However, by proper thermal control of external surfaces, the time required for recooling the magnetto 4.5 K and reenergizing it should be minimized. MAGNET DISCHARGE If any detectable portion of the magnet coil should become normally conducting and not recover to superconducting, the contained energy will be dissipated in an external discharge resistor. This resistor is sized to Iimit both voltage and maximum conductor temperature. A conservative method for estimating the maximum conductor temperature is to compute an adiabatic temperature rise using 0 ['' J, 0 1 2 R dT = ~ Lm; f Tmax To • c; dT (1) Instantaneous thermal communication between the conductor and adjoining materials is assumed. The participating materials are the superconductor, copper, interlayer and interturn insulation, and helium gas. Tbe current history is defined as {Io, I_ - 10 exp(-T/ To), (2) Since R can be expressedas pL/ A, (1) can be restated in the following form of f~To ( A Td To 1) =tm; f +2 Tmax To (c;/p)dT (3) Solution of (3) yields the maximum temperature as a function of delay until a discharge is initiated. Figure 2 shows this curve for a current decay time constant of 69 s. This decay constant corresponds to a peak voltage of 1000 V across a 0.17-.0 quench resistor, with 12 H inductance in the magnet (i.e., To = Hl0 / V max). Copper properties assumed in (3) are those of OFHC copper having a residual resistivity ratio equal to 150 and a 7.76-T magnetic field [2 ]. Figure 2 shows that a maximum temperature of 200 K is reached if a 10-s delay in initiating discharge is allowed. Longer decay time constants (i.e., lower quench voltages) result in higher temperatures. This temperature is considered permissible since it is a conservative estimate and is limited to a small region of the coil where both initial transition to normal conduction and peak field could occur. A 200-K temperature rise from 5 K should result in less than 0.1% thermal expansion of the J. H. V..s..t ud R. M. 312 a- X "'E t-. z Q) ~ Q) a. E J!l E E ::J X "' ~ 0 10 20 40 30 50 Discharge delay time,rd, s Fig. 2. Maximum conductor temperaturein MFIF as a function of delay time. The decay time constant To = 69 s. conductor. Also, 10 s is adequate time to inititate a discharge with an automatic quench detecting system. Magnet structural materials can also develop resistive heating during a fast discharge by a transformer coupfing eflect between the magnet coil and surrounding structure. Because of its properties, the copper guard-vacuum bladder will show the greatest temperature increase from this source. Assuming adiabatic conditions, resistive and sensible heating of the copper can be equated by the relation f. 00 V 2 /R dT 0 = L.,m f Tmax cdT (4) To Fora 1392 : 1 turn ratio, 1000-V peak coil voltage, and a 69-s current decay time constant, this equation becomes 0.26~0 = ( Tmax pc dT YL., (5) JTo Solution of this equation yields a maximum temperature of 70 K for the copper bladder, which is an acceptable temperature increase. A similar analysis of the stainless steel case yields a temperature of 13 K. Total energy dissipation in the case, bladder, and coil jacket is approximately 7 MJ, or less than 1.5% of the energy contained by the magnet before discharge. HELIUM CIRCULADON Important to the thermal control of the magnet is an adequate circulation of liquid helium. Forced pumping is not practical, so natural circulation was chosen because it has been satisfactory in smaller magnet systems. A computer model of the liquid helium natural circulation was developed to estimate steady-state mass ßow 313 Thermal Control for the MFI'F Mqnet Return valve 27m~ I 9m Heliumtransfer lines Fig. 3. Schematic diagram of the liquid helium circulation. The helium transfer Iines are 15-cm schedule 10 pipes. rate and vapor quality. Also, a sensitivity study was made to determine which parameters have the greatest inßuence and to estimate the range of uncertainty for the liquid helium ßow rate. The principal requirement of the liquid helium system is to maintain quality with less than 10 vol.% vapor in the magnet. Heat transfer analyses indicate that significantly higher vapor qualities would probably inhibit cryostability. A schematic of the liquid helium system, shown in Fig. 3, illustrates the lengths, bends, and altitudes of the pipes, the dewar, and the magnet. Note that the magnet inlet and outlet are approximately 16 and 9 m, respectively, below the liquid helium dewar. The conductor pack and magnet shape do not immediately lend themselves to an obvious ßow-modeling approach. Therefore, in modeling the magnet three approaches were considered: the Blasius friction equation, Darcy's porous media equation, and a three-dimensional orifice model In selecting an appropriate method, estimates of hydraulic diameter, ßow tortuosity, porosity, permeability, friction factor, and effective orifice dimensions were made, and the three approaches were compared by means of their respective pressure drops. The porous media approach resulted in the smallest pressure drop, the Blasius approach yielded a pressure drop ten times greater, and the orifice approach gave a pressure drop that was a thousand times greater. The orifice assumption was rejected as unrealistic. The Blasius iriction method was chosen over the porous media approach because the former was more conservative. Thus, the magnet pressure loss was estimated with a modified Blasius friction equation in the following form: e]. 4P = (/LmTld,.M 2 /2-yA;q, 2 ) (6) Flow in the magnet is expected to be laminar by the Reynolds number definition, so that the friction factor is given as f = 64JJ.Apt!J/ 71Md,. (7) 314 I. H. VaaSaat aad R. M. Rllll8 Pressure Iosses in the piping system were simply modeled using loss coefficients for bends, valves, entrance-exit regions, and other effects. These were calculated as functions of the friction factor f. The effect of two-phase ßow was also included by using the Lockhart and Martinelli correlation [5 ]. Flow in the piping system was assumed turbulent. Helium ßow rates were estimated by an iterative computing method. Using an assigned heat Ioad for selected ßow model elements and an assumed ßow rate, helium properlies were determined for each element node using NBS data [6 ]. The resulting pressure imbalance in the ßow circuit owing to cumulative contributions of friction, momentum, and gravity was computed. The ßow rate was readjusted and the calculations repeated until the pressure imbalance was acceptably small. This procedure provided a means for determining the effect of heat Ioad on ßow rate (Fig. 4a). Also, vapor quality was determined and appears as a function of heat Ioad in Fig. 4b. [ 4] --- ---- (b) 200 --300 700 Fig. 4. Computed helium tlow rate (a) and vapor quality (b) as a function of the system heat Ioad. The dashed curves above and below the graphs show the extent of uncertainty in each calculation. (Assumed liquid helium head in the supply dewar is 1m.) Thermal Control for the MFfF Magnet 315 Table 111. Uncertainty Range of Flow Parameters Parameter Range Friction factor, f Tortuosity factor, 1J Porosity, 4J Liquid helium dewar head, m Hydraulic diameter, dh, cm Piping heat leakage, W Length of piping, m 0.014 to 0.018 1.5 to 2.0 0.30 to 0.35 0.5 to 1.0 0.15 to 0.35 30 to 80 80 to 100 Table IV. Liquid Helium System Heat Sources Heat input, w Source Liquid nitrogen shield radiation Liquid nitrogen shield conduction ~agnethangerrods Conductor joints Instrumentation Ieads Helium ducts Total 160 45 60 45 70 80 460 Because the modeling procedure entails some uncertainties, it was of interest to determine how sensitive the results were to changes in certain variables. Table 111 shows the maximum expected range in these flow parameters, and Table IV shows an estimate for the totalliquid helium system heat input. The results of the sensitivity study are reflected in Figs. 4a and 4b by the uncertainty range curves. The effects of two-phase flow are significant for helium at mass qualities as low as only 1% and flows near 300 g/s. Surprisingly, the static head inside the liquid helium dewar is the most significant parameter affecting mass flow rate because of the relatively low fluid ftow resistance of liquid helium. The pipe friction is far more influential on flow than either magnet friction or heat input. On the basis of this calculation a pipe with a 15-cm ID was selected for the liquid helium supply and return lines. This pipe size and the estimated system heat Ioad of 460 W yield an equilibrium mass flow rate of approximately 700 g/s. Vapor quality at the top of the magnet is less than 4 vol.% (0. 7% by mass), and is less than 20 vol.% at the top of the return line. These results imply that an adequate safety margin has been provided in the thermal control of the magnet. NOTATION A = conductor area AP = area for conductor pack c = specific heat dh = hydraulic diameter f = friction factor H = inductance I= current 10 = initial current i = material index J. H. V..S..t ud R. M. a.. 316 L = oonductor length Le = eflective length L". = 8ow path length througb the magnet m = IIUIIII per unit length M = IIUIIII ftow rate R=resistance T = temperature V= voltage Greek Symbols 'Y = density ." = tortuosity factor ,.,. = viscosity p = resistivity -r" = delay time -r0 = current decay time oonstant t/J = porosity REFERENCES 1. R. F. O'Neill and D. H. Riemer, "Thermodynamic Analysis of the Magnet System for Mirror Fusion 2. 3. 4. 5. 6. Test Facility," General Dynamics Convair Division, San Diego, Califomia, CASD-LLL-78-002 (October 1978). Handbook on Materials for Superconducting Machinery, Metals and Ceramies Information Center, Sattelle Memorial Institute, Columbus, Ohio (1974). R. B. Jacobs, "Helium Circulation in the MFTF Magnet.System," Lawrence Livermore Laboratory, Livermore, California, internal oommunication (November 8, 1977). Standard Handbook for Mechanical Engineers, 8th ed. McGraw-Hill Book Company, New York (1978). Fluid Flow Data Book, General Electric Co., Schenectady, New York (1977). R. D. McCarty, NBS Tech. Note 631 (1972). DISCUSSION Question by P. W. Eckels, Westinghouse ElectricCorporation: Have anyofyourexperimentsindicated a maximum permissible vapor fraction in the conductor? Answer by author: No, they have not. However, we believe thatour 10vol.% vapor Iimit isconservative on the basis of allowable heat transfer area decrease without loss of cryostability for the MFTF superconductor. F-7 FORCED-CIRCULATION COOLING SYSTEM FOR THE ARGONNE SUPERCONDUCTING HEAVY-ION LINAC* J. M. Nixon and L. M. BoUinger Argonne National Laboratory Argonne, Il/inois INTRODUCTION The Argonne superconducting heavy-ion linac is a prototype heavy-ion accelerator used to increase the energy of an ion beam from a tandem electrostatic accelerator C· 2 ]. The accelerating elements are split-ring-type resonators with hollow niobium drift tubes mounted in cylindrical housings. The housings are made of explosively bonded niobium clad copper [3 ' 4 ]. The resonators, along with superconducting solenoid focusing magnets are positioned axially in cryostats on support structures which also serve as helium supply andreturn manifolds, as shown in Fig. 1. The resonators and magnets are cooled by a continuous forced-ftow circulation of liquid helium directly from a refrigerator. The cooling system consists of the refrigerator, a 1000-liter dewar with built-in heat exchanger coil, a 46-m3 helium gas storage tank, three distribution boxes with valves, heat exchangers, and transfer line ports, connected by a 20-m-long, liquidnitrogen-shielded, coaxial distribution line. A plan view of the linac cryostats, a buncher cryostat, and the cooling system components is shown in Fig. 2. DESIGN REQUIREMENTS The principal requirement for the cooling system is that it provide continuous refrigeration of the superconducting resonators and magnets at temperatures of s4.9 K, for long periods of time. The unique shape of the split-ringresonator, and its mounting base-down on the support structure, requires continuous forced-ftow circulation of helium through its drift tubes for adequate cooling. Because the resonators and magnets are connected in parallel, ftow must be divided reliably. For simplicity, fixed restrictors are used to divide the ftow, but this requires that the quality of the entering helium be kept constant regardless of the upstream heat load, otherwise the downstream units could receive variable cooling. Uniform ftow division is accomplished by having the helium enter each cryostat as saturated, or slightly subcooled liquid. This is achieved by heat exchange with the return ftow, as described below. * Work performed under the auspices of the U. S. Department of Energy. 317 318 J. M. Nixon and L. M. BoUinger HELIUM LIQUID MITliDGEN AND LIN! 0 10 SCAU . . Fig. 1. End view of beam-line cryostat. The magnet is behind the resonator. A second required feature is the capability of installing or removing any cryostat without interrupting the operation of the remainder. This is accomplished by providing a pair of transfer lines between the distribution boxes and the adjacent cryostats, together with supply, return, and bypass valves for each cryostat. As the third requirement, the system has to supply liquid helium, allow its withdrawal for use in cooling a separate cryostat for testing resonators before they are installed in the linac, and provide for closed-circuit recovery of the boil-off gas during operation of the linac. 319 Forced-Circulation Cooling System for Linac ~ I TANDEM VAULT 0 2 I 3 4 SCALE (meters) 5 HELIUM REFRIGERATOR BUNCHER L INAC ASSEMBLY AREA : ··: 3 COMPRESSORS i... .. j ON FLOOR ABOVE r··, ;__j . ., . . : .; .•. ..0: . .. '·. O ,_ .. , L.. j RESONATOR TEST CRYOSTAT ·. ·. • • ... ••• • • • 0 . ' .. . ~ Fig. 2. Plan view of superconducting linac cryostats and helium-distribution system. CRYOGENIC SYSTEM DESCRIPTION The helium circuit is shown with some typical heat Ioads in the ftow diagram of Fig. 3. Approximate helium conditions at correspondingpoints araund the circuit are given in the pressure-enthalpy diagram of Fig. 4. The scheme used to control the quality of the outgoing stream in this system is similar in some respects to that used in cooling the magnets of the Fermilab Energy Doubler [5 ]. The outgoing stream is held at higher pressure and temperature than the return stream. Pressure is dropped through a J-T expansion at the far end, and the returning, cooler, two-phase stream picks up heat from the outgoing stream. In the energy doubler, the outgoing stream is not allowed to vaporize, whereas in the present system the outgoing stream vaporizes in each cryostat, and is then condensed and subcooled in heat exchangers before entering the next cryostat. The use of two-phase ftow directly from a refrigerator, and improvement of the quality by heat exchange in a dewar that receives the return ftow, are also features found in the cooling system for the large superconducting solenoid for the MINIMAG [6'7] . . expenment The principal component of the cooling system is a CTI Model1400 refrigerator/liquefier, rated at 95 W at 4.6 K with three compressors and liquid nitrogen precooling. The machine has a variable-speed engine and dual charcoal adsorbers with valves and heaters for regeneration while operating. The approximate mass ftow is 6.7 g/s with three compressors. The refrigerator module is located adjacent to the linac and is connected by bayonets to the end of the distribution line. The compressors are located about 50 m away on a thick concrete ftoor above the tandem. In this J. M. Nixon and L. M. BoWn&er 3%0 2 J- T. MOOEL 1400 REFRIG -BEAM - A BUNCHER HEAT LEAKS LINAC CRYOSTATS Fig. 3. Cooling system helium flow diagram. Assumed heat Ioads are shown. remote location, vibrations from the compressors are isolated from the sensitive resonators. The static heat leaks of the components are constant, but the rf power to the resonators is variable. The heat Ioad of each resonator is usually in the range of 2 to 4 W. An average of 3 W is assumed for this example. In operation of the linac, the resonators are separately phased, and power to the various resonators can be optimized to give maximum acceleration for the available refrigeration capacity. 15.40 10 30 12 H, EN T HALPY, 32 J/g Fig. 4. Pressure-enthalpy diagram for the circuit and heat Ioads of Fig. 3 (m = 6.7 g/s). Forced-Cireulation Cooling System for Linac 321 As shown in Figs. 3 and 4, helium leaves the refrigerator at point 2 and enters heat exchanger HX 1 in the 1000-liter dewar. The quality, or gas fraction, of the fluid at point 2 depends upon its enthalpy at point 1, before the refrigerator J-T expansion, and the pressure of the outgoing stream in the distribution line. With three compressors, m = 6.7 g/s, and if P 2 = 185 kPa, then the quality, x 2 = 0.06, and the heat removed in HX 1 to condense and subcool the ftow to 4. 7 K is 17.6 W. Helium then ftows through cryostat D, absorbs 30 W total heat input, resulting in an X4 of 0.22. HX2 in distributionbox D removes 30 W so that conditions at point 5 are almost the same as at point 3, except that the pressure is Iower. Similarly heat is added in cryostat C and removed in HX3 bringing the ftow to point 7. The heat Ioads of cryostat A and the buncher, being much smaller than the previous ones, are taken together in series. The J-T expansion, 9-10, drops the return stream to a pressure slightly above that of the 1000-liter dewar, and Iowers its temperature to give a reasonable temperature difference in the three heat exchangers. The return stream passes through the shell sides of HX2 and HX3, absorbing heat from the outgoing stream. The quality increases at each heat exchanger, until it enters the dewar with x 12 = 0.73. Part of the returning liquid is evaporated by the heat load from HX1, making the final quality, x 13 = 0.88, leaving 12% of the ftow to accumulate as liquid for this case.lf the total heat load of the system is less than the refrigeration capacity, the liquid Ievel in the dewar rises. If the Ioad exceeds the capacity, the liquid Ievel drops and the excess gas is transferred to the storage tank. The 88% of the ftow that is gas returns to the refrigerator as saturated vapor at the temperature and pressure of the dewar, as at point 1311• MECHANICAL DESIGN AND FABRICATION Main Helium Distribution Line The main helium line connects the refrigerator with the 1000-liter dewar, the D, C, and A distribution boxes, and on through the taodem shielding wall to the buncher for a total Iength of about 20m. The center of the line is 2.44 m above the ftoor. A typical cross section is shown in Fig. 5. The innermost tube carries the higher pressure outgoing ftow, while the annulus between it and the next tube carries the reduced pressure return ftow. The helium lines are surrounded by an RETURNING HELIUM G-10 INSULATOR VACUUM 012345 SCALE .cm Fig. 5. Typical cross section of helium distribution line. 3ZZ J. M. Nlxon and L. M. BolliDger annular radiation shield made up of two larger pipes through which liquid nitrogen is fed continuously. About 5liters/hr of liquid nitrogen is supplied at the refrigerator end and exhausts as gas at the buncher end. The sections of line were preassembled in the shop, and field welded with the distribution boxes and dewar. The entire line is welded except for one joint at the corner near the dewar. 'Ibis joint, which has 0-ring seals for the vacuum jacket, double indium seals on the coaxial helium line, and a nickel-gasketed coupling for the liquid nitrogen line, allows the section of line that connects the refrigerator and dewar to be removed if necessary for maintenance of these components. Dewar and Dewar Neck Insert The 1000-liter dewar is a commercial gas-shielded design modified with a larger neck to accept a 150-mrn-diameter insert. A heat exchanger coil consisting of 18.3 m of 15.8-mm-OD x 0.9-mm-wall copper tubing was installed in the bottom of the helium vessel before it was welded shut. The dewar neck insert extends 0.6 m below the top ftange and contains four helium tubes: high pressure to and from the heat exchanger coil, and low pressure into and out ofthedewar. The high-pressure lines are connected to the coil by means of a pair of nickel-gasketed couplings. The heat leak of the neck insert is minimized by bleeding a small ftow of gas up past heat-intercepting copper disks brazed at intervals to the stainless tubes. Distribution Boxes The distribution boxes are liquid-nitrogen-shielded vacuum vessels located adjacent to the heads of the linac cryostats. They contain two transfer line bayonet ports and manual supply, return, and bypass valves in the outgoing stream. The C and D boxes also contain heat exchangers, HX2 and HX3, which have helical coils of the same length and tube size as the dewar coil, mounted in vertical tanks. The A box contains six manual helium valves, an air-operated J-T valve, a cooldown line with a cold check valve, but no heat exchanger. The helium piping is surrounded by annular liquid nitrogen tanks which are connected in series with the liquid nitrogen shields of the line. All of the boxes and the entire line have a common insulating vacuum. The vessels were assembled, welded, and leak tested in the shop, then erected on site and welded to the sections of line. Figure 6 is a photograph showing the three boxes and the partially completed line during construction. All of the TIG welds were made with an argon purge and helium mass spectrometer leak tested when completed. The system has been thermally cycled a number of times since the initial cooldown and there has never been a leak. IDstrumentation The most useful diagnostic parameters are the dewar liquid Ievel and dewar pressure. Outputs from a superconducting Ievel sensor and a pressure transducer are recorded continuously on a strip chart, and are also used to actuate an alarm when high or low Iimits are exceeded. A differential pressure gage with a ± 125-Pa (±0.5 in. of water) range is connected on one side to the dewar, and on the other side to the dewar through a valve in series with a small reference pressure tank. When the valve is closed, the instantaneous dewar pressure is trapped in the reference tank, and the gage then gives an extremely sensitive indication of the rate of dewar pressure Forced-Circulation Cooling System for Linac 323 Fig. 6. View showing distribution boxes and partially completed line during construction. change. This has proven to be indispensable in adjusting the refrigerator and in making heat Ioad measurements. Other gages indicate pressures in the outgoing stream as it enters the distribution boxes, J-T valve, and the low-pressure return sides of HX2 and HX3. Temperatures are determined by germanium resistance thermometers on the helium lines entering and leaving cryostats C and D, entering cryostat A, and leaving the buncher. Copper-Constantan thermocouples are also attached at the same locations to monitor cooldown. Additional thermocouples give temperatures of the liquid nitrogen shields in the distribution boxes and the temperature of the nitrogen exhaust from the shield circuit. Current to a 250-W heater in the dewar is monitored with a digital voltmeter. Liquid helium Ievels in the shell sides of HX2 and HX3 are determined by superconducting Ievel sensors. OPERATING EXPERIENCE The system has been in operation almost continuously since June 1978, with only a few shutdowns for maintenance and for warmups to clear blockages caused by helium contamination. Linac beam acceleration runs of up to three weeks' duration have been completed with no interruptions caused by the cooling system [8 ]. The extreme flexibility of the system has been demonstrated by its ability to operate with the dewar and the distribution line, or with any combination of the linac cryostats. Figure 7 shows schematically the multiplicity of functions it performs. Experiments to determine the best ways to optimize performance under these varying conditions are continuing tobe made. ReHability of the system has been reasonably good. After the usual learning period, there have been only a few unscheduled stoppages. These were caused by 324 J. M. Nixon and L. M. Bollinaer LINAC PORTABLE DEWAR Fig. 7. Schematic of helium-refrigeration and distribution system, which performs the following functions: (1) cools linac; (2) makes liquid; (3) provides liquid inventory; (4) saves gas from precooling linac; and (5) saves gas from miscellaneous use. several instances of helium contamination, and the premature failure of bearings and an 0-ring on the expansion engine. The source of the most serious helium contamination was found tobe a hairline fatigue crack that developed in a copper elbow in a compressor suction line after about 2300 hr Operation. One of the most demanding functions of the refrigeration system is the cooldown of an accelerator section, in which the elementstobe cooled have a mass of about 1100 kg. Initial efforts to accomplish the cooldown by means of the refrigerator alone were not very successful because the ftow rates of warm gas are severely limited and the coaxial design of the distribution line maximizes this problem. Consequently a procedure has been adopted that depends heavily on the use of liquid helium from the 1000-liter dewar. In essence, the procedure is as follows: The resonators are cooled to approximately 80 K by ftowing liquid nitrogen through precooling lines not shown in Fig. 3. Then, helium from the main distribution line is allowed to ftow through the cryostat, and the outgoing warm gas is returned to the refrigerator by means of the warm-gas recovery line shown in Fig. 3. Simultaneously, liquid helium from a portable dewar is transferred through the previously mentioned precooling lines in the cryostat, and the outgoing gas is also fed into the warm-gas recovery line. Finally, when the resonator temperatures have been lowered to :515 K, the outgoing helium can be returned to the 1000-liter dewar through the normal coaxial path. The complete cooldown cycle requires about 24 hr, of which about 18 hr are used for nitrogen precooling and the remaining 6 hr are required to decrease the temperature from 80 to 4.6 K. Since the helium-induced cooling rate considerably exceeds the capacity of the refrigerator, some excess helium gas is generated and this is pumped by the refrigerator compressors into the storage tank. The net usage of liquid during cooldown is typically 400 Iiters. The heat Iosses of various components of the distribution system have been measured with some care by two methods. One approach is to measure the rate of boil-off of helium gas. The second approach is a heat-balance method in which the unknown heat input associated with some Ioad is balanced by means of the known heat input from the heater in the 1000-liter dewar under the condition that the Forced-Cireulation CooUng System for Linac 325 pressure in the dewar (and hence the refrigeration capacity of the system) is constant. The sensitivity and response time of the system are such that it is feasible to measure a change of load of a small fraction of a watt in an hour. CONCLUSIONS All of the design objectives of the cooling system have been accomplished. Forced circulation of liquid helium directly from a refrigerator through a complex circuit to an extended array of superconducting devices has been demonstrated successfully. ACKNOWLEDGMENTS The success of the system is the result of the efforts of many people. In particular, much of the design of the heat exchangers and distribution system is due toP. C. Vander Arend, Cryogenic Consultants, lncorporated, Allentown, Pennsylvania. An exceptional job of fabrication, welding, assembly, and testing of the system was performed by personnel of the Argonne Centtal Shops and Quality Assurance Divisions. REFERENCES 1. L. M. Bollinger, R. Benaroya, B. E. Qiflt, A. H. Jaffey, K. W. Johnson, T. K. Khoe, C. H. Scheibelhut, K. W. Shepard, and T. P. Wangler, in Proc. 1976 Prown Linear Accelerawr Conference, AECL-5677 (1976), p. 95. 2. L. M. Bollinger, IEEE Trans. Nucl. Sei. NS-24:1076 (1977). 3. R. Benaroya, L. M. Bollinger, A. H. Jafley, T. K. Khoe, M. C. Olesen, C. H. Scheibelhut, K. W. Shepard, and W. A. Wesolowski, IEEE Trans. Magn. Mag-13:516 (1977). 4. K. W. Shepard, C. H. Scheibelhut, R. Benaroya, and L. M. Bollinger, IEEE Trans. Nucl. Sei. NS-24:1147 (1977). 5. P. C. Vander Arend, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p.420. 6. M. A. Green, "MINIMAG Experiment, Large Superconducting Solenoid Magnet, The Cryogenic System," Lawrence Berkeley Labaratory Eng. Note M4834, UCID 3754 (June 1975). 7. M. A. Green, in Advances in CryogenicEngineering, Vol. 21, Plenum Press, New York (1975), p. 24. 8. K. W. Shepard, IEEE 'li"ans. Nucl. Sei. NS-16:3659 (1979). DISCUSSION Question by G. H. Morgan, Brookhaven National Laboratory: You mentioned you were concerned about the parallel ftow arrangement. Was this because of the possibility of oscillations? Answer by author: The concern was that if helium entered the parallel ftow-dividing restrictors (lengths of small-diameter tubing) as a two-phase mixture of varying vapor fraction, then the mass ftow metered to each resonator would vary widely, possibly resulting in insufficient cooling. To prevent this, the aim was to cool the stream so that it always entered the restrictors as liquid, making the ftow supplied to each resonator fairly constant. F-8 ENERGY DOUBLER SATELLITE REFRIGERATOR MAGNET COOLING SYSTEM C. Rode, P. Brindza, and D. Ridüed Fermi National Accelerator Laboratory Batavia, Illinois and S. Stoy Cryogenic Consultants, Inc. Allentown, Pennsylvania INTRODUCI'ION The Fermi National Accelerator Laboratory superconducting accelerator (energy doubler) is a 6-km-long magnetring consisting of 1000 magnets ultimately to be cooled by 24 "satellite" refrigerators. At present there are three operating refrigerators and several strings of magnets, the Iongest being 0.15 km, through which a proton beam of 1.25 x 10 13 protonsperpulse has been transported. MAGNET COOLING To coollong superconducting magnet strings a forced-feed system is required. Whether one uses a pump, JT valve, or wet expander is immaterial; what matters is the pressure drops, temperature increases, and heat transfer coefticients to the coils. Subcooled liquid helium (as opposed to two-phase or supercritical gas) has been chosen with two-phase counterftow heat exchange. Subcooled liquid, i.e., liquid 0.3 K below the boiling point, has the highest heat capacity per unit volume as well as the highest heat transfer coefticients. The temperature and pressure distribution for fs of the doubler ring, based on prior magnet test data and calculations, is shown in Fig. 1. The liquid helium is subcooled by a small heat exchanger located in the supply container. lt then reaches equilibrium after the first magnet, point 3. There is a small increase in temperature, 0.05 K, from point 3 to point 4, owing to the two-phase pressure drop in the return stream from point 5 to point 6. The heat generated by the coil located in the single-phase (subcooled liquid) chamber is removed by heat exchange, vaporizing the liquid in the two-phase chamber. The :ftow is controlled by the J-T valve to 326 Energy Doubler SateUite Refrigerator Magnet Cooling System POINT T(K) I 4.90 4.50 4.55 4 .60 4.47 4.42 4.42 4 52 2 3 4 5 6 7 8 ~ 1.8 1.8 1.8 1.8 1.25 1.2 1.2 12 ~ %LIQUJD 14.22 11 .2 0 11.47 II. 75 11.75 27.49 27.99 31.01 100. 100. 100. 100. 96. 13. 10. 0.1 K Super Heat 327 Fig. 1. Details of cooling loop for a string of cryogenic magnets (is of the total ring). maintain point 8 at 0.1 K of superheat. The shield is cooled with two-phase nitrogen with the discharge as 85 K gas. REFRIGERATION SYSTEM A hybrid system which consists of a 4000 to 5000 liters/hr central helium liquefier (CHL) coupled with a small-diameter liquid feedline to 24 satellite refrigerators has been chosen to cool the doubler (Fig. 2). The feedline supplies liquid helium for both the satellite refrigerators and the magnet Iead ftow as weil as liquid nitrogen for the magnet shields. In addition, there are three gas headers; the highand low-pressure helium headers interconnect the six compressor buildings with the 24 refrigerators, while the nitrogen header serves as the gas return to the central nitrogen plant. There is also a helium interconnection gas line between the compressor building and the central helium plant. The satellites act as amplifiers with a gain of 12 by using the enthalpy of the helium supplied by the centralliquefier as liquid, converting it to 4.5-K refrigeration, and returning it as 278.5-K gas. This arrangement combines advantages of a single central facility with those of individual stand-alone units stationed around the ring. The centralliquefiers have the high efficiency associated with large components, but requirements for distribution of cryogenic liquids and electric power to the service buildings are reduced. The likelihood of continued operation in the event of equipment failure is also significantly improved. Table I gives a summary of the total system requirements, consumption, and production specifications, together with power requirements. Table II gives a summary of the magnet heat Ioads. CENTRAL HELIUM LIQUEFIER The central helium liquefier system consists of three large compressors, a helium liquefier, nitrogen liquefier, purification equipment, and storage tanks. The compressors are surplus compressors from an air separation plant. Two of the three have been modified for helium service, while the third will operate on nitrogen. The nitrogen plant will provide precooling for the helium plant as weil as shield cooling 328 C. Rode, P. Brindza, D. Ricbiecl, aad S. Stoy CENTR AL HELIUM LIOUEFIER SATELLITE REFRIGERATORS AT - SERVICE BUILDINGS ( 4 PER SEC TOR) Fig. 2. Layout of refrigeration system. for transfer line and magnets, the nitrogen production being rated at 2550 liters/hr, which will be supplemented by purchased liquid nitrogen. Both the helium and nitrogen plants will produce subcooled fluids at 3 atm for introduction into the feedline at 4.65 and 80.0 K, respectively. The returns from the feedline will then transfer any excess liquid into two storage dewars. These two dewars, which are the only two in the entire system, serve a dual function. They provide system stability and liquid for a few hours of operation if the central liquefiers have to be shut down for minor repair. SATELLIT E REFRIGER ATOR The unit consists of a 35-ft-long heat exchanger column (Fig. 3) a liquid expansion engine, two flow-splitting subcoolers, and a standby gas expansion engine. The unit has four modes of operation. The primary mode, which is used for the energy doubler, is the "satellite mode." In this mode the unit is continuously supplied Table I. Summary of Refrigeration Requirements and Producaion Figures Refrigeration Requirements 1000 Ge V, 35-s cycle 1000 Ge V, dc w Helium Magnet system helium, 4.6 K satellite consumption Total liters/hr 350 =1,548 =1,898 7,480 7,480 w Nitrogen Equivalent liters/hr 480 100 21,600 4,500 Magnet system nitrogen, 80 K Helium transfer line Central He liquefier for 3446 liters/hr of LHe Total for 3446 liters/hr of LHe Total at max. operation w 19,918 19,918 w liters/hr 350 3,096 3,446 Equivalent liters/hr 480 100 2,068 2,648 3,580 21,600 4,500 kW 6,270 2,552 2,540 Input power requirement 24 Satellites Central helium liquefier Nitrogen reliquefier 11,362 Total Production Figures w Helium Satellite refrigerators Central helium liquefier Nitrogen Liquefier liters/hr 23,184 ::=::4,000 2,550 Table II. 4.6-K Refrigeration Loadsand 80-K Nitrogen Requirements for Highest Heat Load Building 1000 Ge V, dc Each w liters/hr w liters/hr 1000 Ge V, 35-s cycle w liters/hr Refrigeration Loads 34 Dipole magnets 34 Dipole ac losses* 11 Quad magnets 11 Quad ac losses* 1 Pair 5000-A leadst Set end boxes 238.0 7.0 13.0 7.0 11.0 10.0 20.0 Totals 1.05 77.0 11.55 14.0 10.0 20.0 14.0 345.0 25.55 Nitrogen Requirements 34 Dipole magnets 11 Quadrupole magnets Totals * 35-s cycle time. t Seven out of 24 buildings. w 22.4 13.5 w 238.0 442.0 77.0 121.0 10.0 20.0 11.55 14.0 25.55 908.0 w 762 138 762 138 900 900 330 C. Rode, P. Briadza, D. Ridüed, ud S. Stoy - - 45FT cold box - -40 -35 berm ~-- - '!t high P"tnure He heoder ~ 0; / --25 - -20 ---15 ---10 Fig. 3. Cross section of satellite refrigerator and cryogenic feed to the superoonducting magnets in the tunnel. 331 Energy Doubler Satellite Refrigentor Mapet Cooling System Table 111. Satellite Refrigerator Parameters Mode Consumption Satellite Refrigerator Liquefier Energy doubler standby 1291iters/hr He (4.48 g/s) 52 liters/hr N 2 84 liters/hr N 2 59 liters/hr N 2 Production Compressor Pin = 1.05 atm Pout = 20 atm Flow = 57.54 g/s 966W 623W 126 liters/hr LHe 490 W plus 26.6 liters/hr LHe (0.92 g/s) with 4.48 g/s liquid helium (plus 0.5 g/s power Iead ftow) from the centralliquefier. This causes an imbalance in the heat exchanger ftow (53.06-g/s supply vs. 57 .54-g/s return) giving a double pinch at 25 and 5 K. The liquid engine expands from 20 atm to 1.8 atm, producing slightly subcooled liquid. The cold end refrigeration comes from three sources: 44% from the heat exchangers ftow imbalance, 48% from the liquid expander, and 8% from the centralliquefier ftow. In the other three modes, liquid nitrogen is used instead of liquid helium. The standby gas engine is now operated at 30 K for these modes, while the liquid engine produces a two-phase Iiquid-gas mixture. The cold box and expanders have been tested, as described in Table 111 and shown schematically in Fig. 4, in the first three modes and have exceeded design in both the liquefier and refrigerator modes and 3 atm LIQUID FEED LI NE 16 atm lWf1 GAS RElUlN TO CENTAAL tELllll UQIEF I ER 4.48g/sec ( ( VAL VE LOCATED AT B-0) COMPRESSOR BUILDING VALVE LOCATED AT ) REFRIGERATOR COLD BOX 4 .48g/sec 20atm 300K !.05atm 278. 5K • • l L---~ SATELLITE MODE LIQUID NITROGEN 4.2g/sec 20atm LIQUID ENG INE 53.34g/sec 300K 81. OK EXCHANGER 57. 54g/sec 282. 5K #2 79.0K 18.0K 6.0K EXCHANGER EXCHANGER EXCHANGER 1!3 #4A #48 26.60K 17.25K REFRIGERATOR. LIQUEFIER & E.D. STANDBY MODE (refrigerator values illustrated) Fig. 4. Satellite refrigerator modes. LOAD 1.2atm 4.52K C. Rode, P. BriDdza, D. Ridlled, aad S. Stoy 332 90% of design in the first attempt in the satellite mode. The energy doubler standby mode is a mixture of refrigeration and liquefaction modes with a trade-off ratio of 5.0 W to 1.0 liters/hr. This mode is designed to cool strings of magnets without the aid of the CHL both durlog initial construction and later durlog failures of the CHL. This modewas used for both the 10- and 25-magnet Al runs. There are many additional mixtures of satellite and refrigeration modes that can be used if the central liquefier is operating at reduced efficiency. FEEDUNE The liquid helium and nitrogen will be fed to the doubler by a 25-section, 6-km-long vacuum-jacketed loop (see Figs. 2 and 5). The loop runs from the central helium liquefier to A4, around the ring to A3, and then back to the centralliquefier. The nitrogen which is used to cool the shield of the magnets also provides the shield for the feedline. The sections are coupled by two rigid, vacuum-jacketed U-tubes, each with a branch tee to feed the local refrigerator. This will permit installation, testing, and cooldown of one section at a time without interfering with the Operation of the restof the system. With the connection of the last service building, A3, back to the central helium liquefier, any section can be taken out of service for repair, if needed, by feeding the return line in reverse. A maximum 4.5-K heat Ioad of 150 W and a maximum 80-K Ioad of 4500 W for the entire line is anticipated. 20 ATM He HE ADER "u" TUBE CONNECTIONS AND DRAW POINTS 3ATM 3ATM LEGEND HELIUM NITROGEN 20 T023 MAGNET STRING .. ""------A------......., Fig. 5. Satellite refrigerator interconnections. -- Energy Doubler Satellite Refrigentor Mapet Cooliog System 333 GAS LINES The satellite gas piping consists of three gas header loops. A 22-cm-diameter low-pressure helium pipe and a 9-cm-diameter low-pressure nitrogen pipe are located on the wall of the tunnel behind the magnet. The helium pipe is the suction line for the compressors as weil as the main magnet relief and manifold for Iead and cooldown ftow. The nitrogen pipe is the collection header for all shield ftow, precooler ftow, and also nitrogen reliefs. The third header is a high-pressure helium pipe which is located on the main-ring-road side of the berm. Two 9-cm gas headers which connect to the centralliquefier are located at A4. The first is a 5 to 16 atm bidirectional helium gas line. Normally it is used as the gas return for the liquid supplied by the centrat liquefier, transmitting gas into the discharge of the centrat liquefier compressors (13 atm). During startup and ED standby mode the line can also supply gas to the 22-cm header. The second header is teed into the 9-cm hitrogen loop and is the main nitrogen return for the nitrogen liquefier. The compressor system is located in the six "zero" buildings with four compressors per building for maximum capacity. The compressors are connected across the two helium headers with all 24 in parallel. The grouping of compressors into a header system totally decouples cold boxes from compressor operation; e.g., all four compressors at BO can be shut down without shutting down any cold boxes. This, of course, involves the loss of ~ of the total capacity. COOLDOWN AND WARMUP If one attempts to cool down long strings of doubler magnets in the normal operating mode it would take several months or might be altogether impossible. The reason for this is that the magnets are heat exchangers and therefore most of the refrigeration that is supplied is heat exchanged with the return line and then vented. Therefore a single-pass cooling of the single phase is used rather than loop ftow, with the two-phase deadheaded. The wave front is very steep and travels through the magnet string much like a step function through a transmission line; i.e., the discharge remains at room temperature during almost the entire cooldown cycle. Cooldown with the central helium liquefier operational is very straightforward. The satellite is adjusted to the liquefier mode producing 126 liters/hr, which is added to the 200 liters/hr from the centrat liquefier (if one is cooling only one service building this might be as high as 2000 liters/hr, stress, pressure drops, and thermoacoustic oscillations permitting). The cooling fluid is run through the single phase of the magnets, returning to the compressor suction by way of the cooldown lines (see Figs. 1 and 5) where it is recompressed to 20 atm. The excess gas is then returned to the discharge of the centralliquefier compressor (13 atm) where it is reliquefied. The wave front is shut off when it reaches one of the cooldown lines and the magnet J-T valve is opened. When it reaches the second cooldown line, the same procedure is repeated, 1000 Iiters of liquid helium aretransferred from the centralliquefier to fill the magnets, the dry engine is turned off, and the satellite is tuned for the satellite mode. If the centrat liquefier is not operational, it is slightly more complicated. The satellite is programmed to produce 20 K gas and run in this mode for about 44 hr with a ftow rate of about 200 liters/hr; it is then reprogrammed to produce 10 K gas for about 8 hr. Owing to the nonlinearities in the heat capacities of metals the 10-K wave will catch up with the 20-K wave at the cooldown line. When the discharge 334 C. Rode, P. Briudza, D. Ridded, IUid S. Stoy temperature reaches 25 K the cooldown lines are closed and the magnet J-T valves are opened. The satellite is reprogrammed to the ED standby mode, and cools the magnet string to 5 K in a few hours; this is known as a transition. After the transition there is a choice of operation; if cooldown is urgent, 1000 Iiters of liquid is obtained from extemal sources. Otherwise the satellite can be used to fill the magnet. The fill time is inversely proportional to the excess refrigeration available, i.e., the excess expander capacity at the service building and the excess compressor capacity in the whole ring. Therefore, fil1 time can vary from 10 hr to a week. This is the mode that was used in the 25-magnet runs. During the filling time, the magnets were energized and the proton beam is tuned to injection energies. F-9 CRYOGENIC SUPPORT SYSTEM FOR AIRBORNE SUPERCONDUCTING GENERATORS P. J. Kerney and P. A. Lessard CTI-Cryogenics Waltham, Massachusetts INTRODUCTION The U. S. Air Force is currently developing electrical generators for a variety of airborne applications requiring high-power and high-voltage capability Cl The airborne environment imposes severe constraints on volume and weight which make a superconducting generator an attractive approach owing to its small size and high power-to-volume ratio. Since superconducting generators are in direct competition with other energy production schemes for these applications, the design of the cryogenic support subsystem strongly influences system potential. A study*, therefore, has been underway to evaluate techniques and system designs for providing the cryogenic cooling to support an airborne superconducting generator system. This paper describes the system constraints, airborne applications, and the resulting cryogenic cooling requirements. The design methodology and trade-off considerations are discussed for the airborne support system, the ground support system, and the cryogenic transport system. Primary emphasis was placed on minimizing the weight of the airborne system and providing the most cost-effective approach for supplying the cryogenic cooling from the wellhead source to the aircraft. SYSTEM CONSTRAINTS A superconducting generator system is considered attractive for airborne applications employing generator configurations with outputs above 10 MW. A summary of these hypothetical applications is included in Table I. Superconducting generator designs are still in their infancy, and the U. S. Air Force is presently sponsoring several exploratory research and development programs in airborne superconducting generators e-'4]. The present generator configuration imposes two important interface constraints on the airborne cryogenic cooling system. First, owing to the generator flow control system, the helium inlet pressure to the generator cannot be less than 1 atm absolute. Second, because the return gas from the generator is utilized to intercept heat leaks in the torque tube and the power Ieads, the gas exiting the generator is essentially at ambient conditions. The specific airborne cryogenic subsystem weight and cooling requirements for the three applications are listed in Table I. *Air Force Flight Dynamics Labaratory Contract No. F33615-78-C-3413. 335 P. J. Kenaey aad P. A. Leisud 336 Table I. Spedfication and Design Goals Application Airborne Cryogenic On-board LHe cooling system weight generator, capacity, target, MW liters/hr lb Airborne operational time, hr 1 10 10 100 10 2 20 20 200 10 3 50 40 500 6 Operational conditions Continuous operation for 20 years Ground cold ready 30-day ftight Operation Ground cold ready 24-hr ftight operation The ground support subsystem also is subject to important constraints. The subsystem must be capable of delivery, storage, and transfer of all necessary cryogens from point of origin to the airborne system with maximum reliability and minimum cost. Specifically, the support system must (1) maintain the generator at a steadystate standby temperature :580 K, (2) be capable of cooling the generator from the standby steady-state temperature to the 4.2 K operational temperaturein a specified time, and (3) provide adequate base storage to satisfy the cryogenic cooling requirements for a specified length of service. Additionallogistics associated with the ground support subsystem typically include a periodic maintenance cycle for each aircraft requiring the system to be warmed to 300 K and subsequently cooled to the standby temperature. SYSTEM COOLING REQUIREMENTS The ground cryogenic cooling requirements are strongly dependent on the thermal characteristics of the particular generator configuration. The cryogenic subsystem must be capable of supplying sufficient refrigeration for transient as weil as steady-state cooling applications. The steady-state requirement is that the generator be maintained in a cold ready condition on the ground at some specified temperature TR, which is less than 80 K. Two types of generator cooldowns are required: a "partial cooldown" from TR to the 4.2-K operational temperature and a "full cooldown" from ambient to TR following normal maintenance of the system. To establish TR and the resultant cooling requirements and the corresponding cryogenic support equipment, it was necessary to model the thermal characteristics of the generator. A simplified analysis was performed to determine the mass ftow rate of gaseous helium at 4.5 K required to maintain the generator at a specified steady-state temperature. An energy balance on the generator yields mHe hin + QL + Or + QR = mrhr + mLhL (1) The continuity equation yields eJ (2) Assuming that QR is constant at 0.125 w and that the ftow split between mL and mr is the same for all Operating points, (1) and (2) can be combined to estimate generator temperatures as a function of helium ftow. The results, plotted in Fig. 1, are based on estimates of the thermal characteristics of a 20-MW generator design with a 77 -K winding temperature [5 ]. Cryogenic Support System for Airborne Superconduding Generators 337 60 50 0: "'~ 40 C> z ö z ..."' .... ;:: "'>- 0 "' 30 20 ~ "' 10 .2 .3 .4 .5 .6 FLOW TO GENERATO R, 9/s Fig. 1. Generator winding temperature vs. 4.5 K helium gas maintenance ftow. To obtain the helium mass ftow requirements during the cooldown operation, the generator, as a first approximation, was modeled utilizing a lumped-parameter analysis. The transient energy equation for the generator is (3) Assuming the generator is all copper and that the ratio of specific heats (Cg/Cp) is constant at one-half its initial value (i.e., at TR) and assuming that the cooldown to 10 K requires 15 min, at which time liquid helium is added, yields @final --= e E>initial -C /C rizHe p ·--tc mg (4) The results of this analysis are plotted in Fig. 2. The mass of the generator was assumed to scale as the square root of the power generated, thus producing the cooldown curves for the 10- and 50-MW generators. The total ground cooling requirements were calculated for each application and are tabulated in Table II. A standby temperature lower than the originally specified 80 K was used because of the large decrease in required cooldown ftow at lower temperatures. HELIUM DELIVERY CONCEPTS Delivery of helium from its source of supply to the base can be accomplished in three general ways: entirely in liquid form, entirely in gaseous form, or some combination of the two. On the basis of cost comparison alone, it can be demonstrated that shipment in liquid form is preferable to shipment in gaseous form. 338 P. J. Kerney and P. A. Lessard 150 2"' 8 50 MW 125 -' 0 0 '-' 8 ... 100 0 20 MW 0: 5 ß 0: ~ X ... 75 -' 10 MW 0 10 20 70 INIT I AL WI NOING TEMP , K Fig. 2. Transient generator winding cooldown vs. initial winding temperature. Cost data obtained from helium suppliers [6 ] indicate that the transportation costs and capital equipment costs associated with liquid cryogen transport are factors of 5.5 and 3.5 less than the corresponding costs for gaseous cryogen transport, respectively. On the basis of this comparison, it was determined that all transportation of the cryogen be in liquid form. GROUND SUPPORT SUBSYSTEM CONCEPTS The primary objectives of the ground support system are to maintain the superconducting generator at a steady standby temperature and to efficiently handle the liquid helium required for the flight operational conditions indicated in Table I. The principal design constraint for the ground support system concepts was cost. A number of system alternatives were studied and compared on the basis of minimum 20-year life cycle cost. The simplest and most direct approach to meeting the ground support cryogenic requirements would be a no-recovery system in which the required helium is transported to the base in liquid form and stored in large dewars. As indicated in Table II, considerable quantities of helium are required to satisfy the airborne and Table II. Airborne and Ground Cooling Requirements Application Airborne operational hrs LHe required for airborne cooling Annual helium req'd for ground cooling (equivalent Iiters) 1 2 3 219,000 hr/yr 8640 hr/30 days 36 hr/day 2,409,000 liters/year 190,000 liters/30 days 1600 liters/1 day 2,500,000 Iiters 1,450,000 Iiters 278,000 Iiters Cryogenic Support System for Airborne Superconducting Generators 339 ground cooling requirements. In this nonrecovery approach, all the helium is vented to the atmosphere and lost. Owing to the uncertainty of the helium supply ['·8 ] and the subsequent impact on the price of helium, this approachwas ascertained to have a prohibitively high life cycle cost. The most cost-effective approach was determined to be a centrat base liquefier with adequate base storage. In this concept, liquid helium is transported to the base, and all helium utilized for ground cooling is recovered and subsequently reliquefied into suitable dewars. Replenishment helium is only required to accommodate the airharne cooling requirements. The main components of the ground support subsystem are the liquefier module with appropriate transfer lines, base storage dewar(s), purifier, recover compressor, and high-pressure gas storage. The boil-off gas from the storage dewar bank is utilized to maintain the generator at the standby temperature. The return gas from the generator is then used for gas shielding of the transfer lines to minimize lasses. In applications where usage is high, the transportation vehicles also act as the dewar bank. An additional cost trade-off in this approach is to consider reducing the base storage requirements for the independent operational capability by supplying a back up generator for the base liquefier. The cooling requirements and base storage system would then be supplemented by the operating base liquefier. AIRBORNE SUBSYSTEM CONCEPTS The most critical portion of the overall cryogenic system for the superconducting generator is the airharne cooling system. Severe constraints on system weight (Table I) and generator interface requirements preclude a variety of approaches that might otherwise be very attractive. For example, on board reclamation of helium boil-off in application #1 would req\}ire a compressor and gas storage bottlesthat would exceed the 100-lb weight goal by approximately an order of magnitude. The requirement of positive pressure helium gas ftow to the generator eliminates any techniques that would take advantage of the decreased pressure and temperature with altitude. The return gas temperature of 300 K precludes any concept that would utilize the available refrigeration in a cold return stream to minimize the airharne systems size and weight. Methods for on board helium reclamation, via a recovery compressor with high-pressure gas storage, or a completely closed-cycle cooler design were likewise discarded owing to the stringent airharne weight specifications. The only viable approach that meets all of the airharne requirements is an open-cycle dewar with provisions for venting the gaseaus helium to the atmosphere. Figure 3 is a graphical representation of weight vs. capacity for commercially available dewars. The data indicate that commercially available dewars would be heavy and that a specially designed spherical aluminum dewar would better satisfy airharne weight restrictions. The supply dewar is a spherical aluminum configuration with appropriate plumbing, valving, and instrumentation. The design will allow initial purging and cooldown of the system to 4.2 K and recovery of all ftash and boil-off gas associated with this process. Filling of the generator/dewar and supply dewar can be accomplished individually or simultaneously in series or parallel. The system is equipped with adequate safeties and controls for in-ftight venting and pressure relief of either dewar. P. J. Keney ud P. A. Lellard • 1200 • DESIGN WEIGHT GOAL SPHERICAL ALUMINUM (REF. 91 COMMERCIAL GAS SHIELDED (MVE, CVI, LINDE) 1000 "' __.. CD 800 .... ~ 600 ------ ---- 400 200 100 200 . 300 400 500 600 CAPACITY, LITERS SPHERICAL ALUMINUM 700 800 Fig. 3. Liquid helium dewar weight (empty) vs. capacity. SUMMARY The basic cryogenic support system design to satisfy the cooling requirements for an airborne superconducting generator includes liquid helium delivery to the base, a central base liquefier, and a spherical aluminum airborne dewar. Table 111 summarizes the basic component sizes for the hypothetical applications. Also included in the table is the annual helium consumption for each application with and without the base helium reclamation. Table 111 also summarizes the 20-year life cycle cost for each application for various base support options. As the tabulation indicates for all system options in every application, the no-helium-recovery concept is less cost effective than helium reclamation. This conclusion was reached even though the cost analysis assumes the price of helium will rise at a rate less than the gross national product. This is considered a conservative assumption since a number of studies C·8 ] indicate that the helium gas available in proven helium-rieb reserves will approach projected demand in the 1990-2000 time frame with a probable subsequent rise in helium costs. Table 111 also shows that providing a backup generator for the base liquefier is more cost effective in all applications. Table m. Component Sizing Snmmary On-board dewarsize, Application Iiters 1 2 3 125 250 300 Base Jiquefier size, litersihr A* Bt C:f: 100 100 200 200 NA 25 * A, no recovery. t B, standby generator for liquefier included. :f: C, totally independent storage. 20-year life cycle cost, 106 1978$ (discounted) Net helium usage per base, 106 scf/yr Liquid storage, 1000 gal A B c A B c A 130 168 3.5 27.5 45.0 NA 120 93 3.5 165 60 0.26 NA 60 0.26 0.06 43 15 3.0 44 8.2 B c 26 28 5.8 6.9 NA 2.2 Cryogenic Support System for Airborue Supercouducting Geuerators 341 CONCLUSIONS A total cryogenic support system to provide cooling for airborne superconducting generators has been analyzed. The most cost-effective system is one providing for liquid cryogen transport to the base, a central base liquefier with a backup generator for ground cooling and reliquefaction, and a lightweight spherical aluminum open-cycle liquid helium dewar system for airborne operational cooling. On-ground helium reclamation is recommended for all applications. ACKNOWLEDGMENTS The work reported in this paperwas sponsored by the U. S. Air Force, Wright-Patterson Airforce Base, Dayton, Ohio, under Contract No. F33615-78-C-3413; Lt. G. Rondash, Program Manager. The support and guidance provided by Maj. G. Puhl of the Flight Dynamics Laboratory and C. Oberly of the Aero Propulsion Laboratory are gratefully acknowledged. J. L. Smith, Jr. of Massachusetts Institute of Technology has served as the superconducting generator consultant for this study. His assistance in defining the generator interface requirements and the generatorcooldown characteristics was invaluable to the program. NOTATION c. = specific heat of generator, Jlg-K CP = specific heat of helium, Jlg-K h;n = enthalpy of gas entering generator, J I g hL = enthalpy of gas exiting through power Ieads, J I g hT = enthalpy of gas exiting through torque tube, Jlg m. = mass of generator, g mHe = mass ftow of 4.5 K gas to generator, gls mL = mass ftow of gas exiting through power Ieads, gls mT = mass ftow of gas exiting through the torque tube, gls QL = totalleak to generator through power Ieads, W QR =total radiation heat leak to generator, W QT = total heat leak to generator through torque tube, W t =time, s tc = cooldown time, s T = temperature, K TR = generator standby steady-state temperature, K (J = T generator - T heliwm K Blina! = 10.0 K- 4.2 K = 5.8 K B;n;t;al = TR- 4.2 K REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. H. Southall and C. Oberly, IEEE Trans. Magn. Mag-15(1):711 (1979). C. E. Oberly, IEEE Trans. Magn. Mag-13:260 (1977). R. D. Blaugher, J. L. McCabria, and J. H. Parker, Jr., IEEE Trans. Magn. Mag-13:755 (1977). B. B. Gamble, T. Keim, and P. A. Rios, "Superconducting Rotor Research," AFAPL-TR-77-68 General Electric Company (November 1977). B. B. Gamble, private communication. P. Hickey, private communication. "Helium, A Public Policy Problem," report by the Helium Study Committee, Board on Mineral and Energy Resources Commission on Natural Resources, National Research Council, National Academy of Sciences, Washington, D.C. (1978). "The Energy Related Applications of Helium and Recommendations Concerning the Management of the Federal Helium Programs," ERDA-13 (April 1975). H. E. Simpkins and R. L. Reed, in Advances in Cryogenic Engineering, Vol. 12, Plenum Press, New York (1967), p. 640. F-10 ECONOMICS OF CRYOGENIC SYSTEMS FOR SUPERCONDUCTING MAGNETS* G. Y. Robinson, Jr. Massachusetts Institute of Technology Cambridge, Massachusetts INTRODUCTION The majority of the cryogenic systems installed for cooling superconducting magnet systems over the past ten years have been aimed at providing a reliable experimental facility for proving the viability of superconductor systems and for advancing the state of the art of superconducting magnets. As a result, the criteria for cryogenic system selection has been that of lowest initial cost and maximum reliability. Systemsnow being installed and those being designed for future installation will be operated on a continuous basis giving impetus to an economic analysis in which the overall operating cost of the cryogenic facility becomes an increasingly important parameter in the design and selection of the system. In addition, conservation of energy through increased efficiency and conservation of heliumwill become significant design criteria. Economics should be considered early in the design phase to permit evaluation of the "trade-offs" and options available in selection of the system and in the preparation of specifications. This study describes some of the options which should be considered and gives insight as to their impact on operating cost. The first section deals with component equipment costs and presents data which can be used by the designer in making a cost trade-off analysis. In the second section, methodology for an operating cost is presented and, finally, examples of typical trade-offs and their effect on Operating costs are reviewed. The major components of the cryogenic system are indicated in Fig. 1 and include, in addition to the superconducting magnet dewar, the refrigerator /liquefier cold box and compressor, helium transfer lines, liquid helium storage, liquid nitrogen storage, gaseaus helium storage, and helium purification equipment. COSTS Costs for the major components of the cryogenic system are presented in this section. The curves shown are the average cost obtained from three suppliers or the actual costs of equipment purchased within the last year and updated to reftect prices in effect for the second quarter of 1979. * Supported in part by the U. S. Department of Energy. 341 343 Economics of Cryogenic Systems for Superconducting Magnets Fig. 1. Cryogenic system. Refrigerat ors The present-da y cost of helium refrigerato r-liquefier s is shown in Fig. 2 and in 1969 and does not differ widely from costs initially reported by Strobridge of the portion upper the in units for here reported Costs updated by him in 1974. in the Costs inflation. to owing er high somewhat e range-ar W 10-k to 1e curve-th lower portion (100 to 1000 W) have been affected by both inflation and lower thermodyn amic efficiency. However, these increases have been counterac ted by economies owing to standardiz ation and productio n of multiple units. As a result the eJ 106 <f1 ... a: ..J ..J 0 0 ,: <f1 0 u 105 I04 L--,--,-,.,rn.r~-.-~~rrrn~--.-r-r 103 102 10 REFRIGERATOR CAPACITY 1../ATTS AT 4 SK Fig. 2. Helium refrigerator cost. G.Y.R....... k. 344 costs are only slightly higher than they were ten years ago. As the requirement for greater numbers of large refrigeraton increases, a reduction in cost can be expected. Over the past ten years, the emphasis on refrigerator development and procurement has been directed towards holding equipment costs to a minimum and increasing reliability. The trend has shifted to oil-lubricated screw compressors and turbo expanden and away from reciprocating equipment mainly to tak:e advantage of the higher reliability experience with rotary equipment. This trend has led to a general decrease in Carnot efficiency rather than an increase as might be anticipated with advancing development. Refrigerator/liquefien produced in the late 1960s were approachlog 20% of Carnot, whereas units presently available are closer to 10%. lt is clear that there is a need for further development to increase efficiency and still maintain the achieved reliability. The costs of equipment shown in Fig. 2 are based on units using liquid nitrogen precooling. The cost of refrigeraton which do not use liquid nitrogen precooling will be approximately 10% higher in the upper range and up to as much as 50% higher in the low range. Transfer Lines Costs of various sizes of helium transfer lines is shown in Fig. 3. The lower curve indicates cost of "off the shelf" or catalog standard vacuum-insulated transfer line sections suitable for small installations. Installations which require the long length of lines with numerous elbows and joints are approximately twice the cost as shown in 10 800 a: ..... "' :z: "' ~ ~600 ....1 ....1 0 0 ~ 0 u 400 ZOO ~ARD SECTIONS 100 lf.! 2 3 4 TRANSFER LINE SIZE, INCHES Fig. 3. Heliumtransfer line cost. 5 Eeonomia of Cryopaic Systems for Supercooductiug Magoets 345 o-' V> 0 u 103 STORAGE DEWARS, LITERS Fig. 4. Liquid storage Dewar cost. the field-erected curve. The cost of liquid-nitrogen-traced lines or concentric helium lines is on the order of 1t to 2 times that of unshielded lines. Liquid Storage Cost of liquid helium and nitrogen storage dewars is shown in Fig. 4. dewars in the range shown are standard sizes which are produced in moderate quantity. The cost of helium dewars increases with capacity to the 0.6 power, as might be expected. Costs have escalated in line with the general cost index over the past few years and can be expected to continue to increase. Gas Storage Cost of helium storage at 1 atm in gas bags and at 18 atm in cylindrical vessels are shown in Fig. 5. The cylindrical vessels are standard ASME coded 250-psi propane storage tanks which have proven adequate and economical because of quantity of production. Their cost ranges from 35t to 50t per pound of steel. Specially designed tanks of nonstandard dimensionswill cost between $0.75 and $1.00 per pound. Helium Purifiers In the larger sizes, helium purifiers are built on a custom basis. Costs are shown in Fig. 6. As the demand for helium recovery purifiers increases it is expected that costs will be reduced slightly. The costs shown are for separate purifier units. Purifiers which are built in as an integral part of the liquefier refrigerator will cost approximately one-third of that shown. OPERATING COSTS Cryogenic system Operating costs include in addition to the direct charges for power, water, and liquid nitrogen, the cost of Iabor and maintenance and a capital G. Y. Robüuoo, Jr. 346 104 o/) a: <( -' -' 0 0 >-' o/) 0 u 10 3 10 2 L----,-,-,-,"<T~---,-,-,-,"<T~---,-,-,3 102 10 10 HELIUM GAS STORAGE, EQUIVALENT LIQUID LITERS Fig. 5. Heliumgasstorage cost. :Q <( -' -' 0 0 ,_. o/) 0 u 10 .I HELIUM PURIFIERS, CUBIC METERS/MIN Fig. 6. Helium purifier cost. Eeonomics of Cryogenic Systems for Superconducting Magnets 347 Table I. Operating Cost Basis Schedule operation Power Water Liquid nitrogen Helium Labor Maintenance Capital charge Installed cost 7700 hr per year $0.035 per kW hr $0.04 per 1000 gal $0.07 per Iiter $0.035 per $7.00 per operating hr 3% of equipment cost 15% of installed cost 2 times equipment cost re Table II. Refrigerator Operating Cost Power Water Liquid nitrogen Labor Maintenance Capital Annual operating cost $53,900 740 32,340 53,900 7,500 74,000 $222,380 charge which is a percentage of the installed equipment cost. In the cost analysis, presented here, the installed cost is assumed to be twice the equipment cost and includes cost of foundations, piping, etc., but not site and building. The basis for the operating cost analysis is shown in Table I. Actual costs will vary depending on location of the facility. However, the model is a reasonable average and provides a basis from which sensitivity of the various factors can be determined. A typical refrigerator producing 200 W of refrigeration at 4.5 K costs $250,000 and requires 200 kW of power and liquid nitrogen precooling at the rate of 60 liters/hr. The annual operating cost of this unit calculated on the basis of TableI is shown in Table II. COST TRADE-OFFS In this section several cases are cited as examples of operating cost analysis with the aim of indicating areas where overall costs can be reduced. Case 1-Refrigerator Location Refrigerators are sometimes regarded as service equipment and located remotely from the superconducting magnet without forethought as to the impact of transfer line heat leak and pressure drop, which often exceeds the refrigeration requirement for the equipment being cooled and sometimes doubles the cost of the refrigerator. Some magnet installations, because of fringe fields, make it necessary to locate the refrigerator a distance of 30 to 50 m from the magnet. In this case, liquidnitrogen-traced transfer lines should be considered. A typical vacuum-insulated untraced transfer line has a heat leak of 0.65 W /m. If the length of transfer line is a total of 100m, then the heat lass will be 65 W, which is much higher than the lass in the magnet dewar, which might be in the 20- to 30-W range. If we assume that a G. Y. a--.., Jr. 200-W refrigerator is required, it will cost $250,000. The cost of unshielded transfer lines will be $40,000 and the annual operating cost is $235,000. Liquid-nitrogentraced transfer lines will cost $80,000 but the heat loss will be only 0.065 W /m or a total of 6.5 W. The refrigerator capacity can be reduced to 140 W at a cost of $200,000 and an annual operating cost of $214,000. The result is an initial investment reduction of $10,000 and an annual savings of $21,000. Case D-He6um and Refrigeradon Recovery Most magnet systems are designed to permit retum of cold gas to the refrigerator with significant savings in investment and operating cost. In this example, consider a cryopump system which requires 45 liters/hr liquid boil-off and the following alternatives: (A) purchase liquid helium and vent boil-off; (B) purchase helium gas, liquefy and vent boil-off; (C) liquefy andrecoverwarm helium; and (D) operate as closed loop with cold gas retum to the refrigerator. At a cost of $2.25 per Iiter of helium the annual cost of case Ais $780,000. The liquefierfor case B costs $150,000 and requires 100 kW ofpower and 45 liters/hrof liquid nitrogen. The annual operating cost is $155,000 and the cost of makeup helium at $3.50 per 100 is $322,000 foratotal of $477,000. For case C the same liquefier is used with an added recovery system and purifier, making the annual operating cost a total of $170,000. A closed-loop cold gas return system will require a refrigerator of 30- to 40-W capacity requiring only one-fourth the power and liquid nitrogen and will have an annual operating cost which is less than $100,000. In summary, the Operating costs would be (A) $780,000, (B) $477,000, (C) $170,000, and (D) $100,000. In addition to the significant cost saving in cases C and D there is also the valuable saving of helium gas. te Case 111-Refrigerator Ef&ciency Improvement As indicated above, the demand for increased reliability in refrigerators has led to the production of equipment which have Carnot efficiencies which are considerably lower than those indicated by Strobridge. This case will show the effect of this lower efficiency on operating cost. Consider the case of the 200-W refrigerator in Table II which costs $250,000. This unit has a Camot efficiency of 6.6%. A refrigerator of similar capacity purchased in 1967 has a Carnot efficiency of 18.3%. If the same equipment cost were assumed, then the annual operating cost of the more efficient unit would be $195,260 with an annual power cost saving of $28,000. Alternatively, for the same annual operating cost, the initial cost could be $343,000. It appears that added investment is warranted in order to reduce power requirements. There are a number of ways in which Carnot efficiency can be increased from the 10% to the 20% range at a moderate increase in equipment cost. These include increase in the number of compressor stages, use of larger or more efficient heat exchangers, increase in the number of expander stages, and introduction of cold recompression. Bach of these steps reduces the loss due to thermodynamic irreversibility in the cycle and has been employed to improve efficiency in a number of cases. For example, the installation of a third expansionstage or "wet" expander in the MIT liquefier in 1969 resulted in an increased capacity of 33%. This in turn, increased the Carnot efficiency from 13.8 to 19.1 %. This marked increase in efficiency was recognized by commercial operators and resulted in subsequent liquefaction plants being retrofitted with an additional expander. Economics of Cryogenic Systems for Superconducting Maguets 349 In other cases it has been shown that by taking the steps indicated above, efficiency can be increased from 11 to 24%. The payoff for development in this area is not only realized in the dollar saving but also in a substantial saving in energy for a given system. ACKNOWLEDGMENTS The author acknowledges with thanks, the support of J. L. Smith, Jr. and the assistance of numerous suppliers of cryogenic equipment in providing cost information, including Intermagnetics General, Cryogenic Technology, Inc., CVI Inc., MVE, Eaton Meta! Products, and Flexiliner Corporation. REFERENCES 1. T. R. Strobridge, NBS Tech. Note 655 (June 1974). 2. F. J. Kadi and R. C. Longsworth, "Assessment and Study of Concepts and Methods of Cryogenic Refrigeration for Superconducting Transmission Cables," Rept. C00-2552-6, U. S. ERDA Contract No. E(11-1)-2552 (February 1976). 3. T. R. Strobridge and R. 0. Voth, IEEE Trans. Nucl. Sei. NS-24(3):1222 (1977). 4. M. A. Green, H. S. Pines, and P. A. Doyle, Cryogenics 19(2):81 (1979). 5. H. Quack and Ch. Trepp, Sulzer Tech. Rev. 60(4):157 (1978). 6. R. W. Johnson, S. C. Collins, and J. L. Smith, Jr., in Advances in Cryogenic Engineering, Vol. 16, Plenum Press, New York (1971), p. 171. DISCUSSION Question by G. H. Morgan, Brookhaven National Laboratory: I am surprised to hear that people are accepting a factor-of-2 increase in refrigerator power consumption in turn for increased reliability. Could you comment on this? Answer by author: First Iet me say that this increase in power required is not entirely owing to reliability requirements. The desire to produce a unitat the lowest possible initial cost is also a factor. With regard to reliability, I believe the approachwas entirely justifiable and that a goal has been reached. I think what is required now is an emphasis on the part of the refrigerator manufacturers toward improving efficiency and a recognition on the part of the users that increased initial cost is warranted in order to reduce power requirements. F-11 MINIMIZATION OF REFRIGERATION POWER FOR LARGE CRYOGENIC SYSTEMS M. A. Hilal Michigan Technological University Houghton, Michigan and Y. M. Eyssa University of Wisconsin-Madison Madison, Wisconsin INTRODUCTION Some cryogenic systems, such as superconducting magnetic energy storage and superconducting generators, require Ioad-hearing supports to transfer forces to a room temperature (warm) structure. lt is necessary to minimize the refrigeration power required to overcome heat leaks through the supports in order to improve system efficiency. In previous studies of heat flow optimization of mechanical supports [ 1-4], it was shown that absolute minimum refrigeration power can be achieved with an infinite series of infinitesimally small refrigerators intercepting the heat leak as it flows from the bot end to the cold end. It was also shown that the absolute minimum can be approached by intercepting the heat leak at a finite number of locations. Bejan and Smith showed that coollog a mechanical SUpport by a variable mass flow rate of helium gas generates the same amount of entropy as an optimized infinite number of refrigerators. They also concluded that the optimum can be approached very closely by cooling with one stream of helium gas, and that continuous cooling of mechanical supports by helium gas is better than discrete coollog by a finite number of independent refrigerators. This conclusion is only true if the gas stream is optimally refrigerated back to the low-end temperature, an arrangement that also requires an infinite series of infinitesimally small refrigerators or an equivalent system to cool the gas stream back to the low-end temperature. In this paper a simple system is considered where the refrigeration power required is mainly due to heat leak through the supports. In this case the cold gas can either be used to cool the supports, thus reducing the heat leak at the ends, or can be retumed to the refrigerator heat exchangers to improve the coeffi.cient of performance. The goal here is to optimize the supports and the refrigerators simultaneously, which, in turn, minimizes the required refrigeration power rather than improving the coeffi.cient of performance of the refrigerator. Optimization results eJ 350 351 Mioimizadon of Refrigeradon Power for Large Cryogenic Systems ~ 0.003- ~ u Fig. 1. Coefficient of performance for a two-engine 1.8-K refrigerator vs. high pressure. 0.002- I ""' 0.0010~~5-_"",:,.---:':-1--.:17-~·=---=I~ 10 15 20 25 30 35 PRESSURE , atm reported in this paper are for cryogenic systems operating at 1.8 or 4.2 K. Only optimization of large-size systems is considered since high-efficiency expansion engines, compressors, and heat exchangers are absolutely necessary in those systems. The supportsarealso optimized using discrete cooling and thus require separate refrigerators operating at different temperatures. In this case the refrigerators are optimized at different temperatures independent of the supports. The optimized refrigerators are used to intercept heat at finite locations on the supports. Supports with variable cross-sectional areas are also considered in this paper. This is done to reduce the refrigerator power if the allowable stress varies with temperature. REFRIGERATOR OPTIMIZATION The Claude cycle with multiple expansion engines and an expansion valve at the cold end is the basis for the present optimization study. Engine arrangements, the T-S diagram, optimization details, and optimization results for 4.2-K refrigerators are published elsewhere [5 ]. That study showed that the cycle is optimized by changing the high pressure and the inlet temperature of each engine. The computerprogram developed in the prior study for 4.2-K refrigerators has been modified and used in the present study to optimize cryogenic systems using superfluid helium. The optimization results for 1.8-K cryogenic systems are reported here. Figure 1 shows the coefficient of performance vs. high pressure for two expansion engine refrigerators operating at 1.8 K. The superfluid helium properties used in the refrigerator program were based on the tabulated values reported by Brooks and Donnelly [6 ]. The optimum inlet temperatures and the optimum mass flow rates through the engines are also given in Table I. In this optimization calculation the compressor efficiency was assumed tobe 85%. The low-temperature Table I. Optimum Parametersfora 1.8-K Two-Engine Refrigerator First engine Second engine Inlet temp., K Exit temp., K Fraction ftow rate 291.959 36.959 62.901 7.896 0.04293 0.12967 M. A.IIIW _. Y. M. E,.a 35% and the high-temperature engine were also assumed to be 85% efficient. The temperature diflerences across the first, second,and third heat exchangers were 10, 0.5, and 0.5 K, respectively. The low pressurewas set equal to the helium vapor pressure at 1.8 K. OPTIMIZATION OF REFRIGERATORS FOR CRYOGENIC SUPPORTS In large superconducting systems such as energy storage magnets and electric generators, most of the need for refrigerationpower is owing to the heat leak through supports. lt was reported [2 ' 3 ] previously that for systems where the heat leak through the mechanical supports dominates the heat Ioad in a cryogenic vessel, it is advantageous to use the evaporating gas to cool the supports. This section of the paper will show that the system is optimized if a fraction of the vapor is used to continuously cool the supports while the rest is retumed to the refrigerator heat exchanger to improve the coefficient of performance. This case is also compared to that of intercepting the heat Ioad at a finite number of locations using separate refrigerators. In both cases, only large refrigerator systems are considered. Supports Coutinuously Cooled by Evaporating Gases Consider the refrigerator system shown in Fig. 2, where a fraction I of the refrigerant is used to cool a mechanical support and the remaining part is retumed to the heat exchangers. The refrigerator is designed to remove an amount of heat, Oe, conducted to the cold end through a mechanical support. This is represented by Oe= Axm (1) where A is the latent heat, x is the liquefied fraction, and m is the total mass ftow rate. Given a mass ftow rate, Im, to cool the support and a heat leak Oe. the length to the cross-sectional area, L/ A, of the support can be determined (see the Appendix). By changing the value of I the required value of L/ A can be determined. For a given fraction f, the refrigerator system is optimized for best coefficient of performance, independent of the support, as reported previously [5 ]. H we change the total mass ftow rate m it is possible to calculate the total refrigeration power as a function of the Support mass ftow rate me =Im. Results of these calculations are shown in Figs. 3 and 4 for epoxy-fiberglass supports with a cold end at 1.8 K and 304 stainless steel support with the cold end at 4.2 K, respectively. It is also shown that the coefficient of performance decreases significantly as more ftow is diverted to cool the support. 1 FIRST ENGINE J SECOND ENGINE xm Fig. 2. Refrigerator-support system in which a fraction of the ftow rate is used to cool the support. 353 Minimization of Refrigeration Power for Luge Cryogenic Systems 3 ...u z <( 'I a: 0 .... <( ... ...J ...a: Q. a: w Q. ...0 ...z ;t 0 Q. ... ~ ......... 0 ...J <( .... u ~ mCp , WATT/K Fig. 3. Total refrigeration power and coefficient of performance vs. heat capacity associated with a mass ftow rate to epoxy-fiberglass supports with a cold-end temperature of 1.8 K. Supports Cooled by a Finite Number of Refrigerated Shields The heat leak to the cold end is intercepted at a finite number of locations using separate refrigerated shields. An optimization procedure reported in a previous publication Cl was used to minimize the total refrigeration power. Using the coefficient of performance optimized for large refrigeration systems, the optimum temperatures, locations, and refrigeration power are calculated for different numbers of shields. The total refrigeration power using any number of shields is less than the refrigeration power required by vapor-cooled supports. Table II shows the optimization results for a 304 stainless steel support with a cold end at 4.2 K and an epoxy-fiberglass support with a cold end at 1.8 K. ':" Q " w u z ~ <l: ~ a: u ~ 0 ~ ...J 6 n. u.. a: w n. 0 .3 mCp. WATT/K Fig. 4. Total refrigerator power and coefficient of performance vs. heat capacity associated with a mass ftow rate to 304 ss supports with a cold-end temperature 4.2 K. 354 M. A. Hllal and Y. M. Eyssa Table II. Optimum Temperatures, Locations, and Refrigeration Power for Finite Number of Shields Compared to Continuous C~oling Number of shields T2,K T~>K T3,K AXdL AX2/L AX3/L AX4/L 304 Stainless steel at Tc = 4. 2 K 0 1 2 3 CC* 39.7 21.6 13.7 81.7 40.7 108.8 0.338 0.189 0.110 0.662 0.334 0.201 0.477 0.295 0.394 Epoxy-fiberglass uniaxial, S-954/E787, in plane Tc = 1.8 K 0 1 2 3 30.8 12.4 7.78 cc 67 .6 30.8 102.5 0.463 0.299 0.223 0.537 0.314 0.232 0.387 0.252 0.293 PL/A, W/cm 5345.0 1060.0 758.4 664.8 4895.0 512.4 112.9 78.6 68.8 406.0 * Continuous cooling. REFRIGERATION POWER FOR VARIABLE-CROSS-SECTION SUPPORTS In many cases the allowable stress on cryogenic supports is a strong fupction of the temperature. The refrigeration power can be reduced using a varial}Je crosssectional area support to take advantage of high stresses at low temperat'ures. Consider the variable-cross-section support shown in Fig. 5. Let th~re be N shields, which intercept heat flowing from the bot end TH = TN+2 to the cold end at T1 = Tc. The cross-sectional area between the two shields is determined by the mechanicalload and the lowest allowable stress in the temperature range of the two shields. At each shield a refrigerator with inputpower Pi will remove heat at the rate llQ = Qi - Qi+ Using calculus of variation, it can be proven that the absolute minimum refrigeration power (N = oo) is Pmin = ~oT1f~H[k(T)/~(T)r /2 dTr (2) where A 0 is the cross-sectional area at TH• L is the support length, k is the local thermal conductivity, and u" is the local design stress normalized to its valut at room temperature. TN+2: TH: 300 K / / / / / / / ' C ! ! f / / / / / / /1/ •o r,., --~ ~ ~ ..~- ~· --: T1 T;., -- _ _ o.c ~•-L ~ _ _ _ _o !:!.-_a,.:!~ 7777777777))77777/7777 T; • Tc L l Fig. 5. Variable cross-sectional area support model. JSS Minimization of Refrigeration Power for Large Cryogenic Systems Table 111. Absolute Minimum Refrigeration Power for Epoxy-Fiberglass Uniaxial, S-994/E787, in Plane C1 300 K) (TH = Absolute minimum refrigeration power Lower-end Temp., Tc PL/A, constant cross section W/cm, variable cross section 1.4 1.8 4.2 25.8 24.0 18.3 18.7 17.5 13.5 Results for a variable-cross-section support are shown in Table III. The results are compared to that öf the constant-area cross section reported by Hilal and Boom C]. For fiberglass-epoxy the refrigeration power using variable crosssectional area supports is reduced by 26% if the cold end temperature is 4.2 K. The shield temperature and location forafinite number of shields can be calculated using equations (16), (17), (18), and (20) in the paper by Hilaland Boom Cl To take into consideration that the allowable stress is a function of temperature, D; CJ. 1 D; = un(T; + 1) [ e;( T~-1)- C;+l(T;~l -1)] t;•+l kdT T T T (3) where un(T; + 1) is the normalized allowable stress at the shield temperature Ti+l· Optimized calculations for epoxy-fiberglass uniaxial, S-994/E787, in plane CJ at Tc of 1.8 K, and 4.2 K for a constant cross-sectional area and a variable cross-sectional area, are shown in Table IV. The coefficients of performance used in these calculations are those optimized for a large refrigeration system. Table IV. Optimum Temperatures, Locations, and Refrigeration Power for Finite Number of Shields (Epoxy-Fiberglass Uniaxial, S-994/E787, in Plane) C1 Nurober of shields T~>K Tz,K T 3 ,K ll.XdL fl.Xz/L ll.X3 /L ll.X4/L PL/A, W/cm 0.293 0.294 112.9 94.08 78.67 61.99 68.85 53.16 Tc= 1.8 K 1 1 2 2 3 3 C* Vt c V c V 30.8 36.16 '12.4 14.23 7.78 8.38 67.6 80.62 30.8 36.3 102.5 116.6 0.463 0.466 0.299 0.303 0.223 0.220 0.537 0.534 0.314 0.316 0.232 0.238 0.387 0.381 0.252 0.248 Tc= 4.2K 1 1 2 2 3 3 c V c c V V 43.6 51.0 21.13 23.65 14.20 15.49 * Constant cross section. t Variable cross section. 85.2 98.4 43.3 51.6 121.5 134.5 0.456 0.463 0.298 0.230 0.219 0.221 0.544 0.537 0.317 0.319 0.228 0.238 0.385 0.381 0.263 0.249 0.290 0.292 72.24 60.55 55.29 44.016 50.016 39.048 356 M. A. Hilalud Y. M. Eyssa DISCUSSION OF RESULTS AND CONCLUSION The present study shows that in certain applications refrigerators should not be optimized independent of the cryogenic system. For a mechanical support system in particular, optimum operating conditions exist where a fraction of the helium vapor is used to cool the support and the remaining part is returned to the refrigerator. The present study also shows that discrete cooling using any number of refrigerated stations requires less refrigeration power than continuous cooling. For discrete cooling the refrigerators can be optimized independently for the best coefficient of performance. In the case of continuous cooling, some of the ftow is used to cool the support and thus affects the performance of the refrigerator. Further reduction in refrigeration power is possible by varying the crosssectional area between shields, taking advantage of the fact that allowable stresses are strong functions of temperature. It can also be concluded that high coefficients of performance for large systems can be obtained. In this case more efficient compressors, heat exchangers and expansion engines can be used. NOTATION = strut cross-sectional area = strut cross-sectional area at room temperature = efficiency constant = mathematical expressiontobe used with equation (18) of reference = force = strut length = temperature =cold-end temperature = hot-end temperature Oe = heat conducted to cold end CP = helium specific heat k = thermal conductivity m = total helium mass ftow rate rilc = helium mass ftow rate to the support s = length variable x = liquefied fraction Un = normalized stress to room temperature value u 0 = allowable stress at room temperature A = heat of vaporization A A0 C D, F L T Tc TH eJ. REFERENCES 1. M. Hila! and R. W. Boom, in Advances in Cryogenic Engineering, Vol. 22, Plenum Press, New York (1977), p. 224. 2. A. Bejan and J. L. Smith, Jr., Cryogenics 14(3):158 (1974). 3. A. Bejan and J. L. Smith, Jr., in Advances in Cryogenic Engineering, Vol. 21, Plenum Press, New York (1976), p. 247. 4. A. Bejan, Cryogenics 15(5):290 (1975). 5. M. Hila!, Cryogenics 19(7):415 (1979). 6. J. Brooks and R. Donnelly, "The Calculated Properties of Helium II," Tech. Report, The Institute of Theoretical Science, University of Oregon, Eugene, Oregon (1973). 7. M. D. Campbell, "Thermophysical Properties of Plastic Materialsand Composites to Liquid Hydrogen Temp. (-423°F)," ML-TDR-64-33, Part III, Air Force Material Laboratory, Wright-Patterson Air Force Base, Ohio (August 1965). Minimizadon of Refrigeradon Power for Luge Cryogenic Systems 357 APPENDIX Consider the support shown in Fig. 3. Assurne perfect heat exchange between the helium gas and the support. For such conditions, d dT dT ds kA ds = rhcCp ds (A-1) Integrating this expression yields dT = Oe ds kA- . + mccp(T - Tc) (A-2) For the support system then F 1 F A=--=-u(T) Uo Un (A-3) Substituting (A-3) into (A-2) and integrating results in L F/uo = I TH Tc --=---k'--10"-:-"=---=:-:Oe+ rhcCp(T- Tc) (A-4) For a fixed L, F, and u 0 there is a relation between Oe and rhccp. For constant cross-sectional area struts, (A-4) reduces to (A-5) DISCUSSION Question by G. H. Morgan, Brookhaven National Laboratory: Since the cross section is determined by mechanical stress requirements, how can the cross section be allowed to vary? Answer by author: The cross section is determined by both the magnetic Ioad and the mechanical design strength. Since the mechanical design strength varies with temperature, it is possible to have different cross sections between shields as shown in Fig. 5. G-1 TWO-DIMENSIONAL HEAT TRANSFER TO SUPERFLUID HELIUM* M. A. Hilal Michigan Technological University Houghton, Michigan INTRODUCTION Steady-state boiling heat transfer to superfluid helium has been studied by many investigators 3 ]. In previous experiments a metalblock was attached to the end of a tube filled with superfluid helium. When the block is heated, internal convection currents take place which result in one-dimensional flow inside the tube. Matthews and Leonard ["] studied boiling heat transfer from reetangular surfaces in a saturated superfluid helium bath. They measured the critical heat flux as a function of orientation, temperature, and depth. In the present experiment the heat transfer from a cylindrical surface immersed in superfluid has been studied attemperatures of 1.9, 1.95, 2.05, and 2.1 K. The flow of both the superfluid helium and normal helium components in this study can be considered as a two-dimensional flow; the critical and the recovery heat flux, two important parameters in the äesign of cryogenic systems are expected to be different from values measured previously 3 ]. The purpose of the present study was to measure both the critical and the recovery heat flux as a function of helium Ievel above the surface. The results of these measurements were then compared with the results of Matthews and Leonard [4 ]. The flow of normal helium leaving the surface can be considered similar to the flow of conventional fluidsernerging from a circular cross section. No attempt in this study has been made to extend the model to cover Kapitza resistance under two-dimensional flow conditions. Rather, the model is used to explain the large increase in the critical heat flux in this two-dimensional flow. c- c- EXPERIMENTAL APPARATUS The various parts of the experiment are shown in Fig. 1. An insulated OFHC copper cylinder, 1. 91 cm in diameter, was suspended in a superfluid helium bath contained in a 11.3-cm-diameter dewar. The helium Ievel inside the dewar was measured using a superconducting liquid Ievel indicator and the helium bath temperature was measured using a high-precision vacuum gauge. The copper cylinder was 1.25 cm high and was surrounded by a nylon block for insulation (Fig. 2). Evanohm wire was wound around the cylinder as a heater. A thermocouple junction, gold-0.07-at.%-iron vs. chromel, was soldered to the * Work conducted at the University of Wisconsin as part of the energy storage project and supported by the U. S. Department of Energy. 358 Two-Dimensional Heat Transfer to Superfluid Helium 359 CU ELEMENT --- --------7'"---~ --== SUPERFLUID HELIUM = Fig. 1. Experimental apparatus. copper cylinder near the surface exposed to the superfluid helium and the reference junction was immersed in the helium bath. Both facing-up and facing-down orientations were used. The power input to the heater was determined by measuring the voltage and the current independently. The calibration tables published by Sparkset al. [5 ] were used to calculate the temperature difference. Fig. 2. Test cylinder. 360 M.A.Hßal , I • lo---+ 0 2.05 K - 3 25 cm MI.." lovel • 2.<Y.> K - 20cm ".,.." lovol c, 2.05 K - 24 .6 cm helium Ievei 9 L9 K - 26.7cm nel;um IOYtl N E ~ .. ö :I: I .5 Fig. 3. Boiling curve for different liquid Ievels and different temperatures. RESULTS The heat ftux vs. temperature difference is shown in Fig. 3 for different temperatures and different helium Ievels. The critical (maximum) heat ftux depends both on the bath temperature and the liquid Ievel. A limited number of data points was taken in the film boiling region. In this region a long time was required to reach steady state and the liquid helium Ievel could have changed appreciably. Figure 4 shows the maximum heat ftux as a function of the helium Ievel at different temperatures. Figure 5 shows the recovery heat ftux also as a function of liquid Ievel at different temperatures and orientations. DISCUSSION OF RESULTS As can be seen from Fig. 3, the critical heat ftux in superfluid helium occurs at large temperature differences. At a temperature of 2.05 K and for 20 cm of liquid Ievel, the temperature difference is 3.4 K. The helium pressure at the surface of the experiment is 3.93 KPa, which is lower than the Iambda pressure. At these large temperature differences the surface can be exposed to either superfluid helium or helium vapor. The critical heat ftux is plotted vs. the liquid head above the surface as shown in Fig. 4. For liquid Ievels up to 25 cm the critical heat ftux does not have a strong dependence on the depth. Also, since large temperature differences occur at the critical heat ftux, the results for 1.95, 2.05, and 2.1 K do not differ appreciably. 361 Two-Dimensional Heat Transfer to Superfluid Helium ~ ~~~ 3 N E u '3= " ;;::" I 2 • I I / I / / / / / / / / / • ,/ 0 . • • 0 • • 0 • 1.95 K- orientation 11 down • 2.05 K- orientotion is up • 2.05 K- or~entaflon is down o 2.10 K- onentot1on is up - 2.05 K- onalyhcol --- Matthews ond Leunord doto ö • I 1 o~----L-----I~0----~----~20~--~----~30~----L---~40 Liquid Level , cm Fig. 4. Critical heat flux vs. depth for different temperatures. Todetermine the critical heat flux analytically as a function of depth we assume that the normal helium velocity leaving the surface is given by 1 (1) Vn CX::--a +bz where Vn is the velocity of the normal helium component, z is the variable height, and a and b are constants. Equation (1) represents the velocity distribution for a flow ernerging from a circular cross section jet [6 ] and is valid for both laminar and turbulent flow. It is necessary to mention that (1) is modified to take into account the finite cross sectional area of the surface. Similar velocity distributions for the superfluid helium component are assumed where the surface is considered a sink. Using the GorterMellink coefficient of friction it was possible to determine the critical heat flux as a function of depth. The analytical results are shown in Fig. 4. The critical heat flux obtained analytically is smaller than the experimental value at low liquid Ievels. The deviation between the analytical and the experimental results can be reduced if it is assumed that superheating takes place near the surface. The analytical modelwill be reported in a future publication. The surface orientation, as observed by previous investigators [4 ], did not have a significant eflect on the critical heat flux results. The recovery heat flux at 2.05 and 1.95 K is shown in Fig. 5. The values shown in these figures are average values since the liquid Ievel changed appreciably in this region. It has been noted that the transition to the film boiling region, as reported in previous studies, is associated with a loud noise and that the noise disappears following transition to the Kapitza conductance region. The data of Matthews and Leonard are shown on Fig. 4 for comparison. CONCLUSIONS High critical heat flux is achieved in two-dimensional flow systems in comparison with one-dimensional flow systems. In two-dimensional systems, the M.A.Hßal 362 3 N E u ...... 31: , u::" V ot;,.o 2 V 0 i o 1.95 K- orientcmon is down " 2.05 K- O<ientotion io up " 2.05 K - O<ientalion io down ::r:: V 20 10 Loquid 30 40 Level , cm Fig. 5. Recovery heat ftux vs. depth for different temperatures. ftow resistance is concentrated near the surface, independent of liquid Ievel. The liquid Ievel increase therefore results in more superheat at the surface but does not change the ftow resistance. This should be considered in the design of large cryogenic systems using superfluid helium as a coolant such as superconducting energy storage magnets. A preliminary analytical study indicates that the critical heat ftux from the surface of a cylinder can be predicted using a conventional velocity distribution for ftow ernerging from circular cross section jets. ACKNOWLEDGMENTS The author wishes to thank M. Steinhoff for his assistance during several phases of this project. Thanks arealso expressed toS. W. V an Seiver for bis very helpful discussions and comments. REFERENCES 1. S. W. Van Sciver, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 340. 2. B. W. Clement and T. H. K. Frederking, in Liquid Helium Technology, Pergarnon Press, Oxford, England (1967), p. 49. 3. G. BonMardion, G. Claudet, andP. Seyfert, inProc. 7thlntem. Cryogenic EngineeringConference, IPC Science & Technology Press, Guildford, England (1979), p. 214. 4. D. W. B. Matthews and A. C. Leonard, in Advances in Cryogenic Engineering, Vol. 19, Plenum Press, New York (1974), p. 417. 5. L. L. Sparksand W. J. Hall, NBS Rept. 9712 (1969). 6. H. Schlichting, Boundary-Layer Theory, McGraw-Hill Book Company, New York (1960), Chaps. IX, XXIV. DISCUSSION Question by R. C. Hendricks, NASA Lewis Research Center: From your ftow model it appears as an axisymmetric jet; could you provide additional information relative to your source term? Answer by author: The ftow model considered is applicable for axisymmetric jets. Equation (1) is valid for both laminarandturbulent jets and the constants a and b can be adjusted to fit the experimental results. G-2 HEAT TRANSFER TO HELIUM-li IN CYLINDRICAL GEOMETRIES S. W. Van Seiver and R. L. Lee University of Wisconsin-Madison Madison, Wisconsin INTRODUCTION Renewed interest in the properties of helium-11 has been inspired by numerous new engineering applications for the fluid 3 ]. Of particular concern is the heat transfer behavior and its effect on the design of stable superconducting magnets. Sturlies of heat transfer to helium-11 in linear systems have investigated such phenomena as peak heat transport, including superfluidity breakdown [4 - 7 ], effects of normal fluid turbulence [8 - 9 ], and time-rlependent heat transfer In general, experiments with cylindrical geometries have been limited to measurements of the peak heat flux from a cylindrical heater or wire in an open bath. On the basis of experience with the linear system, it appears clear that to characterize cylindrical heat transfer, a careful accounting of the heat flow in the adjacent helium is necessary. Ioterest in understanding radial heat transfer to helium-11 is related to work on the Wisconsin Superconductive Energy Storage project, which proposes to use a round cross-section conductor cooled in a bath of helium-11 Cl. The present study is aimed at characterizing the radial heat transfer problem, much as has been achieved for linear systems. e- co-n ]. THEORETICAL BACKGROUND The solution to the problern of heat flow in helium-11 has traditionally been cast in the form of the two fluid hydrodynamic equations with an additional term which takes into account the mutual friction between the two interpenetrating fluids, normal and superfluid. The equations which must be solved are C2 ] 1 I 13 -av. at = --p Vp + SVT- pn A V s - Vn (1) and Ps I -vn 13 +Tin- [ V2Vn +-V(V·vn) 1 ] (2) --SVT+p.Av. Pn Pn 3 where v. and vn are the superandnormal fluid velocities and A is the Gorter-MeHink mutual friction parameter. aV= n --Vp 1 - at P 363 364 S. W. Vu SdYer aad R. L. Lee The steady-state (iJv/iJt = 0) solution to heat ftow through a slit has been published [1 2•13]. Experimentally, it has been demonstrated to be a relatively accurate representation of most one-dimensional heat ftow data C4 ]. For largediameter tubes (d > 1 mm), the solution to (1) and (2) simplifies to an expression for the temperature gradient (_!]_) dT = Ap,. dx S p.ST 3 (3) where Q is the heat ftux density through the helium. A simple method of showing the origin of (3) is to reduce (1) to steady-state conditions (iJv8 / iJt = 0) and assume no temperature-induced pressure gradients, Vp = 0. The latter assumption is equivalent to neglecting normal fluid interactions with boundaries, a mechanism that has been shown to be negligible in large systems (d > 1 mm). Using the expression for the conservation of momentum (psVs + p,.v,.) = 0 (4) and assuming that the heat flow is carried by the normal fluid only Q =pSTv,. (5) an identical expression to (3) results. Furthermore, the expression is not limited to linear heat flow, being applicable to all geometries provided the position dependence of the heat flux is taken into account. Consider a cylindrical heating element of radius r 0 with an applied surface heat ftux, The heat flux density as a function of position coordinate r can be written as Oo. Q(r) = Oo(~) (6) The radial coordinate temperature gradient becomes aT = Ap,. (~) 3 ar S p.ST (rr 0) 3 (7) By analogy with linear experiments, the peak heat ftux density in a cylindrical heat. ftow experiment can be determined by integrating (7). With the boundary conditions on the temperature at infinity set to be that of the bath, the resulting expression Q~ = ( -2 ro JT Ps 2 T1 3S4T3 Ap,. )1/3 dT (8) should predict the peak surface heat ftux, Q~, where T1 is the bath temperature and T2 is the maximum allowable temperature of the helium near the heater surface (2.17 K for p > 5 KPa and determined by the hydrostatic head for pressures below 5 KPa). The goal in the present experiments is to test this expression. EXPERIMENT A schematic. of the experimental apparatus is shown in Fig. 1. A cylindrical heater sample machined from OFHC copper is placed between two insulating disks of fiberglass-reinforced polyester, 140 mm OD and 12.7 mm thick. The copper Heat Transfer to Helium-li in Cylindrical Geometries TO ELECTRONICS t 365 TO HIGH PRESSURE GAS HEAT EXCHANGER HED: P=Ps•r Tii.SK NEEDLE VALVE OEWAR'---.... PRESSURE CYLINOER HEAT TRANSFER SAMPLE Fig. 1. Schematic of experimental apparatus. sample is 13.26 mm OD and has 4.76 mm of its length exposed to the helium, giving a total heat transfer surface area of 198 mm 2 • Spacing between the insulating disks is held constant at 4.76 mm to ensure purely radial heat flow. Temperatures are measured by means of six 1/8-W Allen Bradley resistors, nominal value 75 .n, calibrated against a germanium standard resistor. One of the carbon resistors and the germanium resistor are located inside the copper heater sample. The remaining five resistors are ground to a thickness of 1 mm and placed between the insulatingplates at 7.6, 9.7, 12.7, 25.9, and 66.0 mm from the center of the heater. Each resistor is oriented on a different equally spaced radialline. The present apparatus is designed to allow measurements to be made at saturated vapor pressure as weil as subcpoled conditions. For the pressurized results, a separate cylinder surrounds the experiment and is filled through the needle valve. Data near saturated vapor pressure are acquired by removing the pressure cylinder. In this case, the hydrostatic head of helium is determined using a superconducting Ievel indicator. Heat is applied by passing a current through a wire-wound resistor (room temperature resistance of 180 0) inside the sample. Both step function and continuous heating were studied. The resistance thermometers are measured using a potentiometric conductance bridge for the steady-state values and a chart recorder with a constant current supply for the transient measurements. S. W. Van Sdver and R. L. Lee RESULTS AND DISCUSSION Steady-State Measurements Surface heat transfer data for the present experiment are shown in Fig. 2, which is a plot of heat ftux density vs. temperature difference between the heater and adjacent liquid helium. For comparison, the data have been displayed at both 4.2 and 2.1 K to show differences between normal and superfluid helium for the cylindrical geometry. In general, the two sets of results are qualitatively similar. For both cases, there is a small ä T region associated with nucleate boiling for the normal helium and Kapitza resistance phenomena C5 ] for superfluid helium. Heat transfer coefficients, h, in this region are typically araund 1 W I cm 2 K for both sets of data. Critical heat ftux results are substantially different for the two sets of data. At 2.1 K, with a hydrostatic head of 40 cm, the peak heat ftux occurs at 2.1 W I cm 2 and recovery at 1.9 W lcm 2 , roughly 90% of the peak. The normal helium heat transfer has lower values, with a peak at 0.63 Wlcm2 and recovery at 0.26 W lcm 2 of 40% of the peak. These rather low normal helium peak heat fluxes are most probably caused by the confined geometry of the heat transfer surface. A summary of the experimentally determined steady-state peak heat fluxes in superfluid helium are plotted in Fig. 3 as a function of bath temperature. Unfortunately, the apparatus was not equipped to regulate the hydrostatic head, so there is a slight variation in the helium depth with time. This variation is also shown in the figure by the lower dashed line. The observed peak heat flux data have the characteristic behavior seen in this type of experiment near saturated vapor pressure 6 - 18]. A maximum in the peak heat flux occurs near 1.9 K, with a rapid decrease in value at higher temperatures. For linear heat flow experiments, this behavior can be quantitatively understood in terms of the bulkfluid properties. The goal here was to e 2 0 • 2.1 K 4.2 K 6/:l 0 N 1::.~----~-~~~::.::::.~-=~ 0 00 0 0 0 E ,----------------------0 ?I:'u . 0.5 0 0 0 0 . • 0.2 0 0.1 0.05L-___ L_ _ _ _J__J--~_LLLWL----L--L~~~~~--~ 0.05 0.1 0.2 0.5 I 2 5 10 20 liT. K Fig. 2. Heat transfer and boiling curves for current experiment in He II and He I. Heat Transfer to He6um-ß in Cylindrical Geometries 5 4 N E ~ 3 0 -· • . .. • .• -- --- -- 367 50 E () • - --/ 10 Fig. 3. Experimentally determined peak heat ftuxes in superfluid helium. Solid curve is predicted values for indicated hydrostatic heads using (8). apply the analysis discussed in the previous section to the present case de"lling with a cylindrical geometry. The expression for the peak heat ftux density in superfluid helium confined to a cylindrical geometry is given by (8). Evaluating this expression for the conditions in the current experiment, the predicted peak heat ftux is somewhat higher than the experimental values. The solid curve in Fig. 3 shows these predicted values for the given hydrostatic head. The ratio of the theoretical peak heat ftux, Q~h, to the experimental value O!x is nearly constant, i.e., 0~/ O!x = 1.24 ± 0.07 for all temperatures. It should be noted that similar data [ 16] on cylindrical heaters have been analyzed and these also give experimental peak heat ftuxes below the theoretically predicted value, although the ratio is higher. Several possible explanations for the above behavior can be tested in the present experiment. For example, the Gorter-Mellink parameter, determined from linear experimental data [ 7 ' 13 ], has been used in (7) to determine whether it is independent of geometry by evaluating the data for the temperature distribution in the helium. Figures 4 and 5 show the difference in temperature between the bath and the two resistors nearest the heater surface, at 7.6 mm radius and 9.7 mm, vs. the heat ftux. Within the scatter in the data, these results correlate with a straight line on a log-log plot. Furthermore, both fits appear to have slopes very nearly equal to 3, which agrees with the cubic dependence observed in linear experiments. By using (7), the Gorter-Mellink parameter can be calculated to give values of 120 cm-s/g at 1.95 K, 220 cm-s/g at 2.05 K and 310 cm-s/g at 2.1 K. This is within the range of values measured in linear systems C4 J. At this point, there is no clear explanation for the discrepancy between the predicted and experimental peak heat ftuxes. As noted above, the temperature gradients and Gorter-Mellink parameters are in agreement with values measured in linear systems. However, a few possibilities as to the origin of these results are worthy of mention. The most promising explanation lies in the behavior of the helium near the surface. In the cylindrical geometry, the heat ftux density near the sample is s. w. v. SciY• _.. R. L Lee 2.1 K .. 3.5 4 4.5 5 Fig. 4. Temperature difference between thermometer located nearest the heater (r = 7.6 mm) and that in the bath. amplified, inducing the major temperature excursions to occur there. Therefore, any deviations from the normal mutual friction behavior would occur in a region that is not accessible to bulk fluid measurements. This eflect would be amplified for small diameters, which appears to be in agreement with experiment [16]. Alternately, there may be a fundamental error in the assumptions leading to (8). That the form of (8) is correct within a proportionality factor makes this explanation less probable. Further work will be required in this area to formulate a complete understanding of the results. Fig. 5. Temperature difference between thermometer located second nearest the heater (r = 9.7 mm) and that in the bath. Beat TraDSfer to Helium-D in Cylindrical Geometries 'I 'I - 2 0 ~ • • • '·... 1.5f- 369 1.9 K 2.0 K 2.1 K . • 0 'lo o. "' ., ... 0 .- 10.1 At. sec 10 100 Fig. 6. Increase in peak heat ftux during short-time transient heating. Transient Measurements Two types of transient measurements have been conducted in the present study. The first involves applying a step function heat pulse to the cylindrical sample and measuring the length of time to the onset of film boiling. Similar work has been previously reported for a long linear system [11 ]. The results of the present study are plotted in Fig. 6 as the ratio of the transient to the steady-state heat flux, Q/ Q*. This method of displaying the data shows a temperature-independent form for the effect. As can be seen from the data, there is an observed enhancement of a factor of 2 in peak heat flux for times of the order 250 ms. The understanding of short-time high heat transfer rate experiments has been shown to involve an enthalpy balance within the helium adjacent to the heater. For the long linear system, the rate of enthalpy rise has been shown to correlate with the heat flux. A critical ingredient of this type of analysis is a measure of the transient temperature distribution in the bulk fluid, the second type of transient measurement carried out in the present experiment. Two temperature rise distributions are plotted in Fig. 7 just at the onset of film boiling. As before, the approach to understanding these data is to calculate the extra enthalpy associated with these distributions and compare that with the enhanced heating rate during the transient experiment. The two exponential fits in Fig. 7 are the type used to calculate the enthalpy. As an example, consider the 9.4-W I cm2 data. The total calculated enthalpy rise in the helium is 0.032 J/cm2 • This value is only a few percent of what must be absorbed during the 390 ms of transient heating. Similar results are encountered for the lower heat flux data. Consequently, it appears that the bulk fluid outside the experimental region is participating in the absorption of enthalpy. Only a small temperature rise in the surrounding helium bath would be necessary to account for the energy deposited during transient heating. A problern which remains is how the heat is transmitted from the heater to the bath under no measurable temperature differences. S. W. Vu Seiver aad R. L. Lee 370 Q • 5.9 • 9.4 r-r0 , '\",2 .,.c:J Fig. 7. Temperature rise of thermometers in the heliwn at the onset of film boiling. At 5.9 W/cm2 , flt = 1.4 s and at 9.4 W/cm2 , llt = 0.39 s. Bath temperature is 1.9 K. mm ACKNOWLEDGMENTS The authors would like to thank W. Leake for help in reducing the data. The work was supported by the Department of Energy. NOTATION A = Gorter-Mellink mutual friction parameter p = pressure q = heat flux density r = radial position coordinate Q = heat flux Q* = steady-state peak heat flux Q 0 = surface heat flux S = entropy T = temperature t =time v,. = normal fluid velocity v. = superfluid velocity x = one-dimensional position coordinate p,. = normal fluid density p. = superfluid density REFERENCES 1. R. W. Boomet al., "Wisconsin Superconductive Energy Storage Project, Vol. II," University of Wisconsin (January 1976). 2. S. 0. Hong. P. F. Michaelson, I. N. Sviatoslavsky, W. C. Young, and R. W. Boom, IEEE Trans. Magn. Mq-15:756 (1979). Heat Transfer to HeUum-11 in CyUudrical Geometries 371 3. R. Aymar, G. Claudet, C. Deck, R. Duthil, P. Genevey, C. Leloup, J. C. Lottin, J. Parain, P. Seyfert, A. Torossiam, and B. Turck, IEEE Trans. Magn. Mag-15:542 (1979). 4. C. Linnet and T. H. K. Frederking, J. Low Temp. Phys. 21:447 (1975). 5. G. Bon Mardion, G. aaudet, and P. Seyfert, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1979), p. 214. 6. G. Kraflt, J. Low Temp. Phys. 31:441 (1978). 7. S. W. Van Sciver, Cryogenics 18:521 (1978). 8. W. W. Johnson and M. C. Jones, in Advances in Cryogenic Engineering, Vol 23, Plenum Press, New York (1978), p. 363. 9. S. C. Soloski and T. H. K. Frederking, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1978), p. 222. 10. H. Kobayashi, K. Yasukochi, and K. Tokuyama, in Proc. 6th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1976), p. 307. 11. S. W. Van Sciver, Cryogenics 19:385 (1979). 12. C. J. Gorter and J. H. Mellink, Physica 15:285 (1949). 13. W. F. Vinen, Proc. R. Soc. (London) A240:114 (1954). 14. V. D. Arp, Cryogenics 10:96 (1970). 15. N. S. Snyder, NBS Tech. Note 385 (1969). 16. T. H. K. Frederking and R. L. Haben, Cryogenics 8:32 (1968). 17. A. C. Leonard and M. A. Clermont, in Proc. 4th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1973), p. 301. 18. D. W. B. Mathews and A. C. Leonard, in Advances in Cryogenic Engineering, Vol. 19, Plenum Press, New York (1974), p. 417. DISCUSSION Question by P. Seyfert, Centre d'Etudes Nucleaires de Saclay, France: Were the transient heat transfer experiments performed with the pressure vessel in place? Answer by author: Both the steady-state and transient heat transfer experiments were performed at "near saturated vapor pressure" conditions. These are achieved by removing the pressure vessel and measuring the hydrostatic head of helium. Question by S. Caspi, Lawrence Berkeley Laboratory: How does the liquid Ievel enter into the Gorter-Mellink equation? Answer by author: The liquid Ievel enters equation (8) for the peak heat ftux density, Q~, through the upper Iimit to the integral, T2 • The maximum allowable temperature in the helium is therefore T2 = T1 + I pdT 0 -dp dp where dT/ dp is the slope of the vapor pressure curve of liquid helium. G-3 MAXIMUM AND MINIMUM HEAT FLUX AND TEMPERATURE FLUCTUATION IN FILM-BOILING STATES IN SUPERFLUID HELIUM H. Kobayashi and K. Yasuköchi Nihon University, Tokyo, Japan INTRODUCTION Superfluid liquid helium (helium-11), especially the subcooled helium-II, is said to have special advantages as a coolant of superconducting devices, particularly where there is a possibility ofthermal hazard. First of all, helium-II has better coolinß, characteristics than boiling normal helium, that is, a large peak heat flux Omax C· ] (At Omax• film boiling begins and at Omin. and also a large recovery heat flux Omin the liquid near the heated metal surface reenters the stable non-film-boiling.) For the stabilization of superconducting magnets, it is important that the mechanism of heat transfer to helium-II is understood. Investigations have been made into the behavior of transitions between the non-film-boiling state and the film-boiling state at critical heat fluxes (Omax. Omin) both in the saturated and in the subcooled helium-II, as weil as the cooling stability in the film-boiling state. The mechanism of noisy film boiling which emits an audible high-frequency sound in the saturated helium-II [4 ], also was studied. So far, several investigators have observed the noisy film-boiling process by means of high-speed motion picture techniques [5 ] and microphone techniques [6 ]. In the present research, the frequency of the temperature change was measured with the help of a microthermometer and a clear correlation was determined between the cooling condition and the temperature fluctuation of the metal surface. eJ. EXPERIMENT The experiments were performed with a Niebrome foil element (10 #tm thick) with an exposed area of 6.4 mm 2 • This elementwas attached to a resin substrate into which a carbon resistor bad been embedded as a temperature sensor (see Fig. la). The element was heated by supplying a dc current directly through the foil. This element was in direct contact with about 5 Iiters of liquid helium-II. The bath temperature Tb maintained relatively constant because of the large amount of liquid present. The temperature was measured and controlled with the help of a germanium thermometer. The immersion deptb H was regulated manually by Iifting or pushing the support rod through the vacuum-tight seal with the help of a superconducting continuous Ievel indicator. In order to observe the temperature change of the foil 372 Maximum and Minimum Heat Flux and Tempersture Fluctuation in Superfluid Helium liqui d Su rface T 373 5011 H lnsulator Aquadag Carbon Si deV iew NiCr Foil Top Vlew Mic ro-Thermom"l "r (a) ( b) Fig. 1. Schematic drawing of the test element. Legend: (a) side and top view of the Nichrome foil element ; (b) structure of the microthermometer. surface in the film-boiling state, a microthermometer was fabricated of colloidal Aquadag carbon, about 50 1-'m in length and a few microns in thickness, as shown in Fig. lb. The microthermometer responded tothermal oscillations of up to several kilohertz. For some of the experimental work, the heater element was placed at the innermost part of the reetangular channel. The channel bad an open end with dimensions of 50 x 25 x d mm, where d is the variable channel height. In measurements of the temperature ftuctuation, frequencies were determined after a wait of a few minutes to achieve steady state in the film boiling at each heat ftux. Subcooled helium-II (from TA. to 1.62 K) was produced in adewar vessel similar to that used by Bon Mardion et al. C]. RESULTS AND DISCUSSION Qmax&Dd Qmin The general behavior of the heat transfer in the saturated helium-II must be divided into three regimes depending on the immersion depth H. The position of the boundaries for the regimes depends on Tb [4 ]. The following detailed discussion is applicable for a Tb of 1.91 K, as shown in Fig. 2 although the phenomena were observed at other temperatures as weil. Researchers have already established experimentally and theoretically [8 - 12 ] that Omax of saturated helium-II increases with depth in a manner unlike that of subcooled helium-II. Values of Omax seem tobe comparable with the data of the references if the size effect C0 J and the different sample configurations are taken into account. For H < 8 cm (region I), Omin increases and the hysteresis in the Q vs. a T characteristic becomes smaller as H increases. Here, the film boiling is silent, i.e., produces no audible sound. For 8 < H < 22 cm (region II), Omin becomes successively larger and the hysteresis almost disappears at 22 cm, i.e., Omin coincides with Omax· However, the silent boiling can be made noisy by thermal or mechanical shock. If there has been a transition into noisy film boiling from the silent film-boiling state, the hysteresis is fairly large. For H > 22 cm (region III), film botling always is spontaneously accompanied by sound and the hysteresis becomes undesirably large. Consequently Omin must be subdivided into o:nin for silent film boiling and o::,in for noisy film 374 H. Kobayasbi and K. Yasuköcbi 8 6 N E u 3:4 0 0 2 00 10 H . cm Fig. 2. Critical heat ftuxes as a function of the immersion depth at Tb = 1.91 K in saturated helium-11. Solid points represent data obtained on the thin frost-covered surface. boiling. Noisy film boiling reduces the heat transfer and shifts the second rise of Q vs. ll.T towards higher temperatures. On the other band, in subcooled helium-11 the peak heat ftux is fairly large and it is to be noted that there is almost no hysteresis at least for the thin foil element that was used. The general nature of this hysteresis-free behavior was confirmed with the help of foil samples of stainless steel with larger exposed areas. This observation is different from the hysteresis characteristics obtained in earlier work by Van Seiverfora bulk specimen of aluminum C3 ] . The dependence of the critical heat ftuxes on the bath temperature was measured for both saturated and subcooled helium-11. The results at a liquid height of 20 cm are shown in Fig. 3. 15 0 5 Omax 4.2 K 0 ~~~--~--~--~--~--~__J 1.5 1.6 1.7 1.8 Tb 1.9 2.0 K 2.1 T, Fig. 3. Critical heat ftuxes as a function of the bath temperaturein saturated (dashed line) and subcooled (solid line) helium-11 (H = 20 cm). Maximum and Minimum Heat Flux and Tempersture Fluctuation in Superfluid Helium 375 Temperature Fluctuation In subcooled helium-11, within the temperature range of these experiments, there was no noise in the film-boiling state, and thus, the results below all refer to saturated helium-11. Figures 4a-4c show typical traces of the changes in resistivity of the microthermometer due to temperature ftuctuations in the noisy-film-boiling state for d of oo (without a channel), 2 and 4 mm, respectively. The frequency of the thermal oscillation f as a function of H, with d as parameter, at Tb = 1.91 K and Q = 4.5 W I cm 2 , is shown in Fig. 5. In this case only the frequencies of the fundamental modes were evaluated, although the oscillations have several harmonics. The frequency increases steadily with H . Obviously, there are two basic types of oscillation modes; one for no channel (d = oo, Fig. 4a) and the other for fairly narrow channels (d < 3 mm, Fig. 4b). The modes gradually merge with one another at d = 4 mm, Fig. 4c. In the narrower channel, the thermal oscillation seems tobe related not only to the filmlike Iifting off of liquid helium from the metal surface but also to the blow-off of helium-11 along the channel due to the destruction of superftuidity. Thermal noise observed in a long tube C4 ] also might have some relationship with these results. When the metal surface was covered with a thin film of frost, the frequency increased significantly for a d of infinity tagether with Omax and Omin as shown in Figs. 2 and 5. So far, it has not been possible to find a satisfactory explanation for this observation; also, the results themselves might be somewhat ambiguous. However, the fact that the frequency varies in a manner similar tothat of the critical heat ftuxes has been verified. Similar experiments for d not equal to infinity have not yet been performed. For constant values of H, Tb, and d, f decreases with increasing Q; above a certain value, the sound fades out altogether although oscillations continue up to higher Q ·,,·:i'/\\\vwr~'A'V\\\\\,~,\: 1 . ... .. . (a) (b) Fig. 4. Timevariation of the resistance of the microthermometer at Tb= 1.91 K, and Q = 4.5 W /cm2 • Legend: (a) d = oo, H = 12 cm, (b) d = 2 mm, H = 12 cm, (c) d = 4 mm, H = 17 cm. (2 mV/div, 20msec/div, 0.11-'A.) At d=oo the element is exposed to the helium-II without channel. (c) 376 H. Kobayasbi ud K. Yuukkbi Tb:1 .9 1 K Q:4.5W/cm2 400 .; 300 N J: 200 100 H , Fig. 5. Frequency ofthermal noise vs. immersion depth. cm (Fig. 6). At a heat ftux of Q:;,in• that is, durlog the transition from film boiling to non-film-boiling, the frequency of the noise becomes inaudibly high. The characteristics of the oscillations do not change with the orientation of the heated surface within the liquid nor do Q~in• o:;.in• and Omax· The amplitude of the temperature ftuctuations seems tobe of the order of 1 K. However, it is difficult to measure the absolute value of the amplitude accurately as the sensitivity of the microthermometer depends on temperature. Tb:l.91K H :12cm 400 0 d =00 300 N J: 200 100 4 s 6 Q , 7 6 W/cm2 9 10 Fig. 6. Frequency vs. heat-ftux density. Muimum and MiDbnam Heat F1u and Temperatue F1actutloa Ia Saperbid He6um 377 CONCLUSION In helium-11 the effective cooling stability is larger than that in boiling normal helium. In saturated helium-11, however, the noise of film boiling reduces the cooling effect and results in a lowering of the recovery heat ftux. The high frequency of thermal oscillations in the region of film boiling implies good coollog conditions for the case of saturated helium-11. However, as these present experiments have shown, from a practical point of view, that subcooled helium-11 under atmospheric pressure is to be preferred as the coolant of a superconducting magnet. Also, stabilization of metals by subcooled helium-11 provides safeguards relative tothermal hazards since the cooling effect extends to higher peak heat-ftux densities than in any other coolant. The metals will not exhibit an hysteresis in the heat transfer characteristics and will be stable without ftuctuating noisily in the film boiling region. ACKNOWLEDGMENTS The authors wish to thank T. Ogasawara, L. Boesten, and K. Enomoto for their informative comments and assistance. NOTATION d = channel height f = frequency of temperature fluctuation = immersion depth = heat flux density per unit heated surface area Omas = maximum non-film-boiling heat flux (peak heat flux) H Q Qmin = minimum film-boiling heat flux (recovery heat flux) = recovery heat flux from silent film boiling Q::W, = recovery heat flux from noisy film boiling T = temperature of heated element sudace Tb = bath temperature llT = T- Tb Q~ REFERENCES 1. H. Kobayashi, K. Yasuköchi, and K. Tokuyarna, Proc. 6th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1976), p. 307. 2. B. W. Clement and T. H. K. Frederking, in Pure and Applied Cryogenics, Vol. 6, Pergarnon Press, London (1966), p. 49. 3. G. P. Lemieux and A. C. Leonard, in Advances in Cryogenic Engineering, Vol. 13, Plenum Press, New York (1968), p. 624. 4. A. C. Leonard, Proc. 3rd Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1970), p. 109. 5. F. L. Ebright and R. K. lrey, in Advances in Cryogenic Engineering, Vol. 16, Plenum Press, New York (1971), p. 386. 6. P. Bussieres and A. C. Leonard, in Pure and Applied Cryogenics, Vol. 6, Pergarnon Press, London (1966), p. 61. 7. G. Bon Mardion, G. Claudet, P. Seyfert, and J. Verdier, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 358. 8. A. C. Leonard and E. R. Lady, in Advances in Cryogenic Engineering, Vol. 16, Plenum Press, New York (1971), p. 378. 9. J. S. Goodling and R. K. lrey, in Advances in Cryogenic Engineering, Vol. 14, Plenum Press, New York (1969), p. 159. 10. R. L. Haben, R. A. Madsen, A. C. Leonard, and T. H. K. Frederking, in Advances in Cryogenic Engineering, Vol. 17, Plenum Press, New York (1912), p. 323. 11. T. H. K. Frederking and R. L. Haben, Cryogenics 8::J2 (1968). 12. D. W. B. Mathews and A. C. Leonard, in Advances in Cryogenic Engineering, Vol. 19, Plenum Press, New York (1974), p. 417. 13. S. W. Van Sciver, Cryogenics 18:521 (1978). 14. S. W. Van Seiver and 0. Christianson, Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1978), p. 228. TRANSIENT HEAT TRANSFER IN BOILING HELIUM-I AND SUBCOOLED HELlUM-B P. Seyfert, G. Claudet, and M. J. McCall Centre d'Etudes Nucleaires de Grenoble Grenoble, France and R. Aymar Centre d'Etudes Nucleaires de Fontenay aux Roses Fontenay aux Roses, France INTRODUCfiON One of the problems still remaining in the design of reliable superconducting magnets is the accommodation of unavoidable heat inputs to the system. Resin fracture and conductor movement cause localized Iosses while time-varying magnetic fields affect generallosses throughout the coil system. In the initial stages, the thermal balance between these heat inputs and the heat absorbed by the coolant govems the temperature response of the concemed conductor elements. If the Ievel is reached where current sharing appears, an additional temperature-dependent heat source is switched on. The occurrence of recovery depends on the heat transfer properties of the coolant. The contribution of transient effects on heat transfer may be decisive in this respect and merits careful investigation. Normal boiling helium has already been investigated in this regard [1- 5 ]. A few authors have reported on technologically interesting transient heat transfer characteristics of superfluid helium [6-8]. This paper presents a first account of experimental results obtained in the course of a still ongoing program. A similar investigation but using a different experimental technique will be published elsewhere r]. EXPERIMENTAL TECHNIQUE Several test cells were fabricated using a common design of two straight concentric 0.1-mm-thick stainless steel tubes with helium being maintained in the annular space between the two tubes. A 1-mm-thick copper tube carrying a thermocouple and heater coil replaced the centrat section of the outer tube and served as the heated surface. Alltest cells bad an OD of 10 mm and the same heated surface element. Figure 1 shows a cross-sectional view of the compound heating element and the central part of a test cell. 378 Transient Heat Transfer in BoUing HeUnm-1 and Sabcooled HeUnm-D 379 Fig. 1. Cross-sectional view of central part of a test cell. In each of the three versions of electrical heater used, the electrical insulation was achieved by a thin (6-~m) coating of alumina. The coating was obtained by anodizing a 12-~m-thick aluminum layer deposited with an electroplating technique on the copper surface. The processing was carried out by an industriallaboratory*. During experiments it turned out that test cells equipped with all three heater versions gave about the same temperature response, namely, an exponential-like rise of 4 to 5 ms duration. The temperature of the heated surface was sensed by a Au/Fe Chromel thermocouple. A relatively remote position (see detail on Fig. 1) was chosen for the thermocouple junction so as not to disturb the uniformity of the heat flux entering the liquid helium. The thermocouple signal was used as a detector for the transient quasiequilibrium states rather than as an absolute thermometer. Based on the construction principle described above, four different geometries of test cells were devised with the aim of simulating specific situations actually occurring in superconducting magnet windings. Figure 2 shows a schematic drawing of all test cells used. Test cells of type 1 are representative of events giving rise to a disturbance which concerns a large conductor volume. Initially heat released in any conductor segment can only be transferred to the fluid adjacent tothat part of the conductor. On the other band, test cell 4 is representative of a situation initiated by a localized disturbance. All the fluid contained in the affected cooling channel and bath is potentially available for heat transfer. The test cells of type 2 and 3 represent intermediate cases which could contribute to the understanding of the present transfer mechanisms. Fig. 2. Test cells used in study. Types 1, 2, and 3 are virtually closed. Both ends of type 4 are connected to the bath via intermediate constrictions. *Siemens AG, Erlangen, Germany. 2 hole~ 0 3ITvn 310 P. Seyfert, G. Clndet, M. J. McCall, ud R. Ayawo For experiment the test cells were placed in a vacuum container immersed in a bathofliquid helium at atmo5pheric pressure which could be operated either at 4.2 K or in the range of 1.6 to 2.16 K C'1. The closed cells, types 1 to 3, were connected to the bath by fine capillaries which ensured their initial filling and a recovery time of some minutes after the discharge of a pulse. PROCEDURE Initially, the behavior of all test cells subjected to simple reetangular heat pulse disturbances was investigated. In the case of the closed cells any significant excursion in the film boiling regime was avoided because of the dsk of damage due to the resulting overpressure. Since vapor formationwas not expected to cause any harm to the open cell, recovery from transient film boiling was studied exclusively with this cell by applying dual-heat pulses composed of an initial intense reetangular pulse simulating a disturbance, followed by a second pulse of reduced amplitude (designated as postheating) representing the joule heating within a conductor element. Trace records of a typical monopulse experiment are shown in Fig. 3. Heater current traces are Iabeted A, B and corresponding temperature traces are Iabeted a, b. Upon onset of heater power the temperature is seen to reach, in a few milliseconds, a plateau in the range of 0.3 to 2 K above the initial value. If a certain duration of the heat pulse is exceeded (trace B, b) a runaway of the temperature occurs. At this point a limiting energy can be defined as the amount of heat that the liquid has absorbed prior tothermal runaway. lt is believed that the limiting energy is an interesting property of transient heat transfer when applications are concemed. The purpose of the monopulse experiments therefore was to measure this property as a function of heat ftux amplitude. When the dual heat pulses were applied to the open test cell (type 4), trace records as shown in Fig. 4 were obtained. Three temperature traces, Iabeted a, b, c are shown. For better clarity of presentation the corresponding heater current traces are omitted. Thermal runaway is initiated almost immediately by the fi.rst pulse, but a more or less pronounced transient recovery is attained with power reduction to the postheating Ievel. Trace a exhibits a recovery to a temperature difference of the order of 1 K, stays at this Ievel for a while, then rises rapidly to a second Ievel of about 4.2 K and finally rises to values weil above 4.2 K. Increasing the duration of the fi.rst heat pulse reduces tbe duration of the lower recovery Ievel. Trace c always remains above it. Trace b barely reaches this Ievel and corresponds to a limiting situation for recovery. The purpose of the dual-pulseexperimentswas to determine tbe limiting values for initial pulse energy as a function of the postheating power. Fig. 3. Trace records of a typical monopulse experiment. Heater current traces are labeled A. B and c:orresponding temperature response trac:es a, b. Vertical scale for temperature is 1.70 K/div; time scale is 6 ms/div. Transient Heat Transfer in Bolling Helium· I and Subcooled Helium· II 381 o ' I c Fig. 4. Trace record of a typical dual-pulse experiment. Temperature response traces are Iabeted a, b, c. Vertical scale is 3.4 K/div; time scale is 600 ms/div. The initial peaks of traces b, c are off the chart and are at 20 and 23 K, respectively. ' I DISCUSSION OF EXPERIMENTAL RESULTS The mono pulse results are shown in two diagrams which present the previously defined limiting energy vs. heat flux amplitude with both quantities being referred to the unit area of heated surface. Figure 5 shows the results obtained with pressurized superfluid helium contained in test cells of type 1 (1 mm and 2 mm) and with normal helium in the type-4 test cell. Each of the straight lines passing through the origin is associated with a particular pulse duration which is indicated as a parameter. The experimental values have been compared to the enthalpy increase [11 ] of the helium mass for the limiting case of a uniform heating up toT>.. = 2.163 K. The calculated values are indicated as dash-dotted horizontal lines along the ordinate. Up to a flux density of 4 W /cm 2 , the experimental values are between 5 to 8% below this theoreticallimit. Even at 8 W /cm 2 , the experimental values show a negative deviation less than 20% from this Iimit. The first important conclusion that can be drawn from these results is that as long as the propagation distance for heat in pressurized superfluid helium is in the mm range, the fluid participates essentially as a whole for transient heat transfer. As a practical consequence, subcooled helium-11 can act as a thermal bailast by adjusting 0 6 7 8 H.ot flux,W;cm2 Fig. 5. Monopulse experiments. Results obtained with three type-1 test cells (• and x independent cells with same nominal d = 2 mm, [:J d = 1 mm) containing subcooled superfluid helium and one type-4 test cell (V d = 2 mm) containing normal boiling helium. Solid Iines through the data points show trend of the data. 382 P. Seyfert, G. Claudet, M. J. McCall, and R. Aymar the depth of the cooling channels to the estimated heat release of the generalized pulsing disturbances. The behavior of boiling helium-1, on the other hand, is quite different in this respect. Figure 5 shows results obtained in this study together with data from similar measurements published in the literature. Although exhibiting almost the same shape, the curves differ from each other in absolute value by factors of 3 to 5. Part of this discrepancy is probably due to the different sample geometries used. Normalliquidhelium has an extremely low thermal conductivity (10 4 less than copper) which Iimits the depth of the fluid layer that can actively participate in a transient heat transfer event. lwasa Cl observed that increasing the channel gap of his sample arrangement from 0.12 mm to higher values caused the limiting absorbed energy to approach saturation at a gap width of 0.5 mm. Hence, one of the test cells with a channel gap of 2 mm should have benefitted from the best possible values of limiting energy available in normal helium. Considering the present experimental results from this point of view, it is concluded that in normal helium at 4.2 K only a fixed amount of energy (10 mJ/cm 2 ) can be absorbed with heat flux densities below 0.5 to 0.6 W /cm 2 • Finally, monopulse experiments have been performed with the more extended test cell types 2, 3, and 4. Figure 6 shows the results obtained in subcooled helium-11 at 1.8 K. The curves clearly demonstrate the influence of fluid volumes situated at increasing distances from the heated surface as a function of the heat flux. The results obtained with dual-pulse experiments performed in the locally heated open test cell, type 4, are shown in Fig. 7. One should note that unique values for the steady-state Iimit were found for helium-11 at all bath temperatures. Hysteresis effects are apparently absent with this fluid for the test cell geometry used. On the contrary, the well-known hysteresis effect in normal boiling helium-1 gave rise to the two characteristic quantities qpb (occurrence of film boiling with slowly increasing heat flux) and q, (recovery from film boiling with slowly decreasing heat flux). lt appears as a general result for helium-11 and helium-1 that, provided the postheating amplitude does not exceed the steady-state Iimit (q, in the case of helium-1), recovery is secured irrespective of the energy of the initial disturbance ~ 300 c"' ~>.. 200 ~ ~ w 100 10m~ Sm~ 4 6 7 6 H.a t f lu x,W/cm2 Fig. 6. Monopulse experiments. Results obtained with test cells of type 2( +), 3(~) . and 4( Y) containing unsaturated superfluid helium. Transient Heat Transfer in Boiling HeHum-1 and Subcooled HeHum-11 383 T. ::1.60K 1.0 '\5 Po~t-hoaling heol f1u",W/cm2 Fig. 7. Dual-pulse experiments. Results obtained with type-4 test cell. Power ratio of initial pulse to postheating part shown as: • = 10; x = 6; + = 3; 6 = 2; D = 1.5. Vertical straight lines indicate steady-state Iimiting heat ftux values. (this corresponds to full recovery current when referring to stability criteria). In helium-11, recovery is still possible for postheating amplitudes above the steady-state Iimit, if the energy of the disturbance is restricted (criterion of transient stability). As an example, in the particular case of the test cell geometry used at 1.8 K and with a postheating amplitude of 1.5 W /cm 2 (twice the steady-state Iimit) transient recovery occurs despite disturbances as large as 35 mJ/cm 2 • In boiling helium-I a new singular value of heat ftux density qtr appears midway between q, and qpb· No recovery appeared possible above this value. Between q, and q1., there is a zone of conditional recovery, indicated with a hatched stripe in Fig. 7. In fact, the admissible energy values within this zone are those of the limiting energy found in the monopulse experiments (Fig. 6) and consequently depend on the heat ftux of the initial disturbance. lt is evident from the data that considerable benefit is obtained from transient effects in unsaturated helium-11 compared to boiling helium-1. The results of conductor stability experiments [12 ] clearly confirm these findings. The results obtained on pulsive heat transfer by Kobayashi and Yasuköchi [8 ] also indicate an increased cooling capacity of helium-11 compared to helium-1. They show, however, marked differences with the present data which are believed to stem from different experimental conditions: The authors have investigated saturated helium-11, whereas this work considered subcooled helium-11. In addition, their heater pulse sequence is not equivalent to the dual pulse utilized in this study. An additional point concerning subcooled helium-11 should be noted. This study shows that a heat ftux above the steady-state Iimit together with a superimposed energy of an initialpulse can be maintained for a specified time without impairing the recovery capability. Apparently the fluid can remove more heat than the limiting steady-state ftux would predict. A rough estimate supports the assumption that this excess heat is responsible for the temperature profile in the channel. 384 P. Seyfert, G. Claudet, M. J. McC.U, ud R. Aymar CONCLUSIONS The present investigation of transient heat transfer properlies of subcooled superfluid helium compared to those of normal boiling helium was carried out with special attention to potential applications. Experimental conditions which are directly related to problems of superconducting magnet technology have been emphasized. The results of the monopulse experiments provide a means to cope with large disturbances or ac losses. Subcooled superfluid helium responds to heat pulses up to several watts per square centimeter by an initial quasiequilibrium state during which the fluid within a few millimeters from the heated surface absorbs nearly its total energy-carrying capacity up to the Iambda temperature. The associated temperature rise of the heated surface remains almost constant and is moderate in size. Actual values are currently being measured for copper with surface quality similar to conductor matrix material. With an initial bath temperature of 1.8 K for instance, every layer of fluid can accommodate nearly 28 mJ per unit surface area and per millimeter thickness. The corresponding phenomenon in normal boiling helium only gives access to fixed amounts of energy (10 mJ/cm 2) if the heat flux remains below 0.6 W/cm 2 • Accommodation of intense localized disturbances followed by an excursion into the current-sharing state of the concerned conductor segment is simulated with the dual-pulse experiments. When the transient quasiequilibrium state in the superfluid helium is exceeded in a locally heated cooling channel the large momentary temperature excursion is followed by a transient recovery even if the initial pulse is continued with a postheating amplitude above the steady-state limiting flux. This behavior which is not observed in normal boiling helium should allow considerable improvement of stability in magnets cooled with superfluid helium. On the other band, if a permanent heat flux is suddenly switched on in the normal boiling helium, a new limiting heat flux value below the well-known peak nucleate boiling flux appears which cannot be exceeded without jumping into the film boiling region. Results already show that transient heat transfer in subcooled superfluid helium considerably increases the overall heat transfer properties above the steady-state performance. The corresponding contribution of normal boiling helium appears quite small in this respect. The decisive advantage of subcooled superfluid helium is based on its large exploitable enthalpy reservoir. REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. Y. lwasa, M. J. Leupold, and J. E. C. Williams, IEEE Trans. Magn. Mag-13(1):20 (1977). 0. Tsukamoto and S. Kobayashi, J. Appl. Phys. 46(3):1359 (1975). Y. Iwasa and B. A. Apgar, Cryogenics 18(5):267 (1978). W. G. Steward, Intern. J. Heat Mass Transfer 21:863 (1978). C. Schmidt, Appl. Phys. Lett. 3%(12):827 (1978). H. Kobayashi, K. Yasuköchi, and K. Tokuyama, in Proc. 6th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1977), p. 301. S. W. Van Seiver and 0. Christianson, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1979), p. 228; also Cryogenics 18(9):521 (1978). H. Kobayashi and K. Yasukochi, Cryogenics 19(2):93 (1979). D. Gentile and W. V. Massenzahl, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 385. G. Bon Mardion, G. Claudet, P. Seyfert, and J. Verdier, in Advances in Cryogenic Engineering, Vol. 23, Plenum Press, New York (1978), p. 358. J. Maynard, Phys. Rev. B 14(9):3868 (1976). G. Claudet, C. Meuris, J. Parain, and"B. Turck, IEEE Trans. Magn. Mag-15(1):340 (1979). G-5 HEAT TRANSFER MEASUREMENT WITH A SMALL SUPERCONDUCTING COIL SUBJECTED TO TRANSIENT AND QUASISTATIC HEATING AT TEMPERATURES BETWEEN 1.8 AND 4.2 K D. Gentile Centre d'Etudes Nucleaires de Saclay Saclay, France and W. V. Hassenzahl Los Alamos Scientific Labaratory Los Alamos, New Mexico INTRODUCTION This paper reports results of some heat transfer experiments with a small, noninductively wound coil constructed of a monofilamentary, copper-stabilized, NbTi superconductor. These experiments have been made in helium at different temperatures, particularly at 1.8 and 4.2 K. The 1.8-K measurements were performed at the equilibrium vapor pressure and atmospheric pressure. The different tests were in the quasistatic and transient heat transfer regimes. To determine the surface temperature and calculate heat transfer coefficients in helium, it was necessary to measure certain physical characteristics of the conductor. These preliminary tests are described briefty in the first section and references are given to publications describing each experiment. The method of measuring temperatures using monofilamentary conductor was described elsewhere [ 1- 3 ]. EXPERIMENTAL PROCEDURE Apparatus Two identical, noninductively wound coils of a monofilamentary, copperstabilized, NbTi conductor on a 10-mm-diameter, 14-mm-long stainless steel mandrei were used for the measurements. The radial dimensions of the conductor are superconductor, 36-#-1-m diameter, copper, 55-#-1-m diameter, and insulation, 65-#-1-m diameter. For the heat transfer tests the insulation was removed; but, to have good insulation between turns, a 65-#-1-m-diameter nylon monofilamentary spacer 385 D. Geadle aad W. V.IIMieiiZÜI was wound between turns of the superconductor. Each sample bad 101 turns wound from 3.20 m of conductor. The conductor and insulation of one coil were glued to the mandrel; the other bad only a thin insulating layer between the coil and mandrel. Significant ditferences were detected between the heat transfer characteristics of the two coils. Before discussing the results, it is worth noting immediately that the surfaces of these samples were quite irregular with an area of about 2.45 cm2 • If, instead of considering all of the exposed surface, only the surface of the coil covered by the conductor is considered, the area is reduced to about 1.56 cm2 • Heat transfer between a surface and a fluid is normally ~ven as a function of the heated surface. For these calculations a surface of 2.00 cm was used. For transient measurements, the etfective surface is probably closer to 2.45 cm2 , andin the film boiling regime for steady-state conditions it is certainly closer to 1.56 cm2 • Measurement Method The purpose of these experiments is to determine the heat transfer mechanism between a solid and helium. This measurement is accomplished by generating a heat ftux in the sample and obtaining a heat exchange coefficient based on the surface temperature. This method, which uses the superconductor as both sensor and heater, was developed at Saclay by Schmidt [4 ]. It is based on the fact that current in a superconductor decreases linearly with increasing temperature above Tc and that at constant temperature the longitudinal electric field in the superconductor increases exponentially with the transport current eJ. The physical constants of the conductor' such as critical temperature, critical magnetic field, magnetoresistivity of copper and current/voltage characteristics of the conductor must be determined precisely to apply this method. These characteristics are described in a separate report [6 ]. A previous publication has indicated that the systematic error due to the knowledge of the physical constants was about 0.1 K. Other errors in calculating temperature are due to the inaccuracies in the method of measuring voltages across the coil. The maximum error is estimated to be less than 0.5 K. EXPERIMENTAL RESULTS Qnasistatic Beat Transfer Regime This section describes the steady-state heat transfer for three different conditions: 4.2 K, saturated bath; 1.6 K, saturated bath; and 1.8 K, atmospheric pressure. An electric method is used to obtain the heat ftux, q, vs. the temperature ditference, llT = Ts- TB, where Ts is the surface temperature and TB is the bath temperature. From the voltage/current characteristic of the conductor in a constant background magnetic field, one can deduce the heat transfer characteristic from the Joule heat generated in the sample and dissipated in the bath, q = UI/ S, where U is the voltage across the coil, I is the transport current, and S is the wetted surface (==2 cm2 ). The surface temperature relation was developed earlier as C] Tc- TB ( E E) Ts = TB + Ia, + ß ln(E/ Eo) Ia, + ß In Eo -I + RN (1) where Ts is the surface temperature, TB is the bath temperature, Tc is the critical temperature, Eis the longitudinal electric field in the conductor, Ico is the critical Heat Transfer Measurement with a Small Superconducting Coil 387 1o r--------+--------~--------r-~ i ......... GOOOLN>An> I<EY,l : 1.8(12K,CO!'PER - - - GOOOLN3An) IREY, T:1.55 K, COPPER T:l8K - .../ I i ... ......_ B: 3l,T: 1,621(,c:a.c..tated witl>~ :a.3634AIK(oor~· I - ·- T: , .ZK(our r051Jts) T: 1.8K--i'------- -- + - - - -- --=ic-t?-if---i X ~ 1)_, r--------+---------t;-:__,.;LfYL--ir"'"~ !;t ILI :r 2 3 l>T, K Fig. 1. Steady-state heat transfer characteristics at three different temperatures. current corresponding to an electric field E 0 of 1 JL VI cm, ß is the change in current in the conductor when the field changes by one decade (tll = 0.016A), I is the transport current in the conductor, and RN is the normal resistance of the copper in parallel with the superconductor. Figure 1 shows the results of these measurements and also those of Goodling and Irey CJ at 3 T. The curves are very similar; nevertheless, note that the curve for the new data increases more quickly for q > 0.1 WI cm2 • Supposing that the exchanged flux is q = h(tlT)", the data in Fig. 1 give n = 1.06 and h = 0.3 W I cm2 K. During the current rise, a weak heating was noted in the helium bath at 1.6 K. At the beginning, the bath temperature was 1.6 K, corresponding to a transport current of 3 A and just before the transition to film boiling, it was 1. 7 K for a current of 5 A. The 4.2-K curve is nearlyidentical to the 1.6-K curve below a flux of 0.1 W lcm 2 , but there is a significant difference in the peak nucleate boiling maximum. At 4.2 K the peak nucleate boiling maximum is about 0.8 W lcm 2 ; at 1.6 K it is about 4 Wlcm 2 • Other measurements were made at T8 = 1.8 K under atmospheric pressure [8 ]. A maximum heat flux was found before the transition of about 9 WI cm2 , which is much higher than the previously measured values for a temperature difference of about 2.5 K. This curve gives a heat transfer coefficient of h = lim qT· = 0.89Wicm2 -K 4T-O/l (2) 388 D. GentUe and W. V. Hassenzahl soo...w so...w I- 1'\. V '- u- Fig. 2. Voltage across and current in the sample during a 40-ms pulse. Sms Transient Heat Transfer in an Infinite Bath In a transient heat transfer study, heat pulses of 10 f-'S to 10 ms were produced in the sample and the voltage recorded across the sample and across a series shunt resistor. These two voltages, which are shown in Fig. 2, determine Ieo and I for use in (1). Figure 3 shows the conductor surface temperature calculated with (1) as a function of time for various values of q. The temperature remains nearly constant for heat ftuxes below the steady-state maximum ftux, about 10 W /cm 2 • Above this ftux the temperature is constant for a time, ft. and then rises quickly and passes the critical temperature, Tc, above which (1) is no Ionger valid. 25.7 7 21 .3W/cm1 W/cm2 6 • 5 3 .zs.lJ', ---•- -+-+-+r Wfaril 2 o~.~--------~I.~ 0 ----------10~--~ t ,m5 Fig. 3. Conductor surface temperature vs. time, with heat ftux as a parameter. Bath temperature is atl.SK. 389 Heat Transfer Measurement with a Small Superconduc:ting Coil Table I. Coefficients for Heat Transfer in a 1.8-K, 1-Atm LiquidHelium Bath q, a, W/cm 2 Ts, K W/cm 2 K 4 Ts-TB TB hKfh<]. hK 0.57 1.07 2.9 7.87 16 17.3 19 21.3 25.7 30 37 44 2.56 2.98 3.75 4.35 4.9 5 5.51 5.66 6.05 6.4 6.8 7 0.018 0.016 0.0155 0.023 0.028 0.0282 0.021 0.021 0.0194 0.018 0.0174 0.0185 0.422 0.656 1.083 1.417 1.722 1.778 2.061 2.144 2.361 2.556 2.778 2.889 1.83 2.485 4.115 5.845 7.825 8.233 10.53 11.28 13.41 15.54 18.24 19.71 0.75 0.91 1.48 3.08 5.16 5.4 5.12 5.52 6.04 6.52 7.4 8.46 The time constants of the coil and conductor, in particular for the heat transport in the NbTi, are on the order of a few microseconds. The classicallaw that the heat flux is proportional to the difference T1 - T~ has been verified. If Tn remains constant and Ts is taken at about 1 ms on the curves of Fig. 3, q may be expressedas q = a(T1- T~) (3) Table I gives the values for a. This result and the Kapitza theory may be compared. The Kapitza conductance, hK, is given by !!..T 2 3 !!..T [ . hK=_!!_=4aT11+-(-)+(-) Tn 2 Tn !!..T 1 !!..T +-(-) 4 Tn 3 ] (4) where 4aT1 = hCJc = lim.1r-+o (q/ !!..T), and limH-+O (hK! hCJc) ~ 1. For Ts > TA, hK/hCJc increases quickly, which is probably caused by a small helium-1 layer at the solid surface. As heat ftux is increased, the He-1/He-11 interface is not weil defined, and finally apart of the heat ftux, above a certain value qmax. vaporizes the He-l film. Because gas is a good insulation, the removal of the heat from the surface is not very good and the temperature difference increases quickly. This flux, qmax• referred to by 4 12 10 8.5 6.7 4.3 4W/crn 2 2.9W/cm 2 ~:JJJL ~--· '.s Fig. 4. Conductor surface temperature vs. time, with heat flux as a parameter. Bath temperature is at 4.2 K. 390 D. Gentile and W. V. Hassenzahl /I yJ 1.5 :.::_ 1.0 Cl> 0.5 v- 0 4 8 I 12 Q. w/cm 2 Pannetier et al. [ 11 ] is between 1 and 10 Wlcm 2 and is a function of temperature, pressure, geometry, orientation, and immersion depth of the sample. For comparison, similar results are presented at 4.2 K. Figure 4 shows the temperature difference between the conductor and the helium bath vs. time. Note that the temperature is nearly constant for a time and then rises very quickly to the critical temperature. This rapid temperature increase is due to the onset of film boiling. The initial stable value of Ts is given in Fig. 5 and is proportional to q. The heat transfer can thus be described by a single heat transfer coefficient for all q values, h = qiT; = 6 Wlcm 2 -K, which is in agreement with the Kapitza conductance. The classical law of thermal conduction that the heat ftux, q, follows a t- 112 dependence has also been verified. The solution to the classical heat equation gives a relation between q and t of q = TrLD 112 t- 112 , where L is the heat of vaporization (2.4 Jlcm 3 ) and Dis the thermal diffusivity (5 x 10-4 cm 2 ls). This formula gives a q that is a factor 2 above the present experimental results. This agreement is satisfactory because the theory corresponds to exchange on a ftat surface. The difference is probably a result of geometric factors. To sum up, the total energy transferred before onset of film boiling is q = TrLD 112 I t 112 - 0.2 J I cm 2 for t < 1 ms. The heat transfer coefficient is about 6Wicm 2 -K for transient and 0.5Wicm2 -K for steady-state conditions. Transient Heat Transfer in Helium Channels Similar experiments were made with the same coils cooled by He-Il in 1- and 2-mm-thick, concentric channels. To avoid damaging the coil or the helium container, which could occur if the energy deposited were greater than the maximum that could be evacuated, the power deposited in the sample was first fixed for a very short pulse and then the pulse duration was increased until a transition occurred in the sample. Figures 6a and 6b show the evolution of the conductor temperature vs. time. The voltage observed across the coil is nearly constant foratimethat depends on the power, then it rises very quickly to another plateau that indicates the critical temperature has been exceeded. After this plateau is reached, the dissipated power still remains nearly constant because the current is automatically decreased as the resistance and voltage increase. For power dissipation larger than about 9 W I cm2 , the plateaus in voltage and temperature are almost nonexistent and the temperature rise comes very quickly after the beginning of the pulse. The curves of Figs. 6a and 6b indicate that there may be two different heat transfer mechanisms, depending on the value of the heat ftux. Below 9 W lcm 2 the temperature is constant until the transition; above about 9 W lcm 2 the temperature Heat Transfer Measurement with a Small Superconductina Coil 391 8 ~ 6 ...: 5 50 60 10 (a) w---------~~ 3 ~-------~=-------~ t . ms (b) Fig. 6. Evolution of the conductor temperature vs. time for (a) 1-mm and (b) 2-mm channels. Bath temperature is at 1.8 K and pressure is 1 atm. of the helium near the conductor increases immediately to the critical temperature. Also, the transfer is better with a larger channel. One Iimit on the energy that can be accepted by the helium in the channel is fixed by the total enthalpy available between the bath temperature and the transition temperature. After this amount of energy has been deposited in the helium, the heat transfer should decrease very quickly. This energy is f 2 . 17K Eo = Vch 1.8 K pCdT (5) where p is the density, Cis the specific heat, and Vch is the volume of the channel. Figure 7 is a comparison of the Iimit, E 0 , and the amount of energy that was 2 deposited in the helium before a transition occurred. For fluxes above 9 W /cm the product of the power and time before the break in the voltage curve of Fig. 2 was taken as the deposited energy. CONCLUSION S The method of measurement used here has the advantage of a relatively fast response time and makes possible the study of heat transfer with high fluxes with 392 D. Gentile and W. V. Hauenzahl 0 UJ ~0.5 5 10 0 .w/cm 2 Fig. 7. Normalized energy that can be absorbed in a channel vs. generated power. neither separate beater nor separate tbermometer. From tbe experimental results using tbis metbod it appears tbat a quasistatic beat ftux of about 1 W/cm2 can be maintained in 4.2 K liquid belium, wbereas 10 W/cm 2 can be sustained at 1.8 K under a pressure of 1 atm. Experiments in a belium batb at 4.2 and 1.8 K sbow tbat a ftux bigber tban tbe steady-state beat ftux can be maintained for a period tbat decreases as tbe beat ftux increases. Tbe product of pulse lengtb and pulse power decreases witb increasing power. A new pbenomenon bas been observed in tbe experiments witb belium cbannels. Tbere is probably a cbange in tbe beat-transfer mecbanism wben tbe beat ftux into tbe cbannel is bigber tban tbe static beat-ftux Iimit in an infinite batb. To explain tbis pbenomenon, it will be necessary to investigate tbe propagation of beat in He-II. ACKNOWLEDGMENTS The authors wish to thank G. Claudet and P. Seyfert of the CEN/Grenoble for their suggestions and auistance during these experiments. REFERENCES 1. W. Hauenzahl, D. Gentile, and M. Polak, "A Method of Determining Temperatures and Heat Transfer Coefficients with a Superconductive Sample," tobe published in J. Appl. Phys. 2. D. Gentile, W. Hauenzahl, and M. Polak, "Heating of Monofilamentary NbTi Superconductors in tbe Current Sbaring State," CEN/Saclay Rept. Supra 78-73 (1978). 3. D. Gentile, W. Hassenzabl, and M. Polak, Cryogenics 20(1):37 (1980). 4. C. Scbmidt, Appl. Phys. Len. 32(12):827 (1978). 5. M. Polak, I. Hlasnik, and L. Krempasky, Cryogenics 13(12):702 (1973). 6. D. Gentile, W. Hauenzahl, and M. Polak, "Cbaracteristiques d'un monofilament de NbTi a temp6ratures inferieures a 4.2 K," CEN/Saclay Rept. Supra 79-15 (1979). 7. J. S. Goodling and R. K. lrey, in Advances in Cryogenic Engineering, Vol. 14, Plenum Press, New York (1969), p. 159. 8. D. Gentile and W. Hassenzabl, "Exchanges tbermiques entre un echantillon supraconducteur et l'belium superfluide sous pression atmospberique," CEN/Saclay Rept. Supra 79-02 (1979). 9. L. J. Cballis, private communication. 10. D. Gentile, "Etude du transfert thermique entre un echantillon supraconducteur et l'belium en regime transitoire," CEN/Saclay Rept. Supra 78-43 (1978). 11. B. Pannetier, J. Phys. (Paris), supplement au No. 10, Colloque C4 (1972), p. C4. 12. 0. Tsukamoto and S. Kobayashi, Jpn . J. Appl. Phys. 46:1359 (1975). G-6 HEAT TRANSFER OF HELIUM IN A PIPE WITH SUCTION L. L. Vasiliev, G. I. Bobrova, and L. A. Stasevieh The Luikov Heat and Mass Transfer Institute Minsk, BSSR, U.S.S.R. The development of effective means of cooling helium is presently of great significance. The enhancement of such heat transfer and, particularly that due to suction of the fluid through a permeable channel wall has been the subject of many publications. The available Iiterature shows that such sturlies of heat transfer have involved fluid flowing in different geometry channels to determine whether it is possible to enhance this heat transfer process by permitting suction of some of the fluid through the permeable channel wall. Most of these sturlies involve theoretical analyses of the process. The emphasis has been on laminar flow. This is presumably because the theoretical analysis of laminar flow is simpler and more tractable than turbulent flow. The number of experimental studies, however, is limited. Equations applicable for motion, continuity, and energy were used to analyze the steady-state laminar heat transfer. It was shown C· 2 ] about thirty years ago that the motion equation for fully developed laminar flow in tubes with suction or injection through a permeable channel wall may be reduced to a nonlinear differential equation of the fourth order. These investigations were later extended, supplemented, and published 4 ]. Since suction enhances heat transfer in a porous channel flow, previous investigators have focused their attention not only on the solution of the fluid dynamics problern but also on the effect of suction on the temperature distribution along the porous pipe radius and length. Thus, Yuan et al. [3 ] have studied the effect of low injection velocities of the fluid through the porous wall on the wall temperature distribution. Temperature profiles have been calculated by Pederson et al. [5 ] for a flow moving inside a porous pipe at a constant wall temperature. Baithby [6 ], on the other band, studied the fluid dynamics and heat transfer for a constant wall temperature and heat flux; an analysis was made of the effect of injection, suction, and channel geometry on temperature and velocity profiles. It is readily apparent that a comprehensive study of heat transfer of a fluid flowing in permeable pipes, i.e., with suction and injection, involves a great deal of difficulties; understanding of this problern even for laminar flow is still incomplete at this time. For turbulent heat transfer the problern is much more complicated. In this case both suction and injection strongly affect fluid dynamics and heat transfer. The effect of surface suction of heat and mass transfer characteristics for turbulent pipe flow in pipes has been studied analytically by Kinney et al. Cl. These authors have shown e· 393 394 L. L. Vuiliev, G. I. Bobrova, aad L. A. Sauevieh Fig. 1. Experimental arrangement. that suction exerts considerable influence on the Nusselt number, friction coefficients, and velocity profiles. At present, porous heat exchangers find wide use in electric and cryogenic devices. Their design requires a knowledge of the heat transfer coefficient. Therefore, it is important to experimentally verify the analytical relationships that describe turbulent heat transfer to rarefied gases in pipes subjected to suction and injection. There are few studies that verify the theoretical predictions with experimental results. The knowledge in this field also is insufficient to soggest some simple and reliable methods for estimating heat transfer efficiency in porous heat exchangers. This present study is designed to provide additional experimental data on the effect of suction on the heat transfer of helium flowing inside a porous ceramic pipe. Figure 1 shows the experimental equipment; it consists of a traditional horizontal cryostat described elsewhere [8 ] . This device consists of a 5-m-long cryostat (2), pump (6), power supply (1), rotameter (7), gas meters (8), and recording unit (9). The test volume consisted of a 650-mm-long thin-walled stainless pipe, 40 mm in diameter (Fig. 2). The tubewas surrounded with a nitrogen screen (1). A vacuum between the tube and screen served as insulation; the residual gas pressure in the vacuum space was maintained at about 10-5 mm Hg. The nitrogen screen and test volume were placed inside of a 120-mm-diameter pipe (3), which was maintained at room temperature and thermally insulated from the nitrogen bath by means of vacuum insulation. A porous electrically heated tube (5) was located along the axis of the test volume. An adiabatic screen (4) placed 5 mm from the external pipe surface was kept at a temperature between 1 and 2 K, the same as the porous wall temperature. To compensate for temperature deformation, all tubes were fitted with end bushings. Helium at moderate excess pressure, supplied from a 40-liter dewar (3) (Fig. 1), was firsttransferred through a thermally insulated region to achieve hydrodynamic stabilization of the flow before being transferred through the porous tube. As the fluid moved down the porous tube, some of the fluid entered the peripheral channel through the permeable wall. The flow in the centrat and peripheral channels was regulated by valves (4) (see Fig. 1). Gas flow rates were regulated by PC-5 rotameters (7) and then measured by PI-40 gas counters (8) placed in series with the rotameters. From the gas counters the heliumwas returned to the cooling system. The rotameters were used mainly to regulate the flow, i.e., a deviation of the indicating float from the prescribed Ievel indicated a change in the fluid discharge at the inlet of the porous tube. Preliminary measurements of the pressure drop along the length of the porous tube indicated valuesrangingfrom 0.01 to 0.1 atm. Therefore, to simplify the experimental device and measuring system, pressure was monitored in the centrat and peripheral channels using standard MO-type manometers and having a scale division of 0.005 atm (Fig. 2). Heat Transfer of Helium in a Pipe with Sudion 395 Fig. 2. Detail of porous tube. The porous tubes were fabricated by sintering spherical bronze powder (0.2 to 0.315-mm-diameter particles). The porosity of the tube wall varied from 28% to 40% ; the permeability coefficient was equal to 7 x 10-8 cm 2 , the maximum pore size was 190 ~m, while the mean pore size was 60 ~m. The tubewas 600 mm long with an inner diameter of 12 mm and an outer diameter of 17 mm. Tobe passive section with 1/d = 300 was also 12 mm in diameter. Heat was supplied to the tube by direct current from a BY 12/600-type rectifier. The temperature was measured with copper-copper plus iron thermocouples (6) (Fig. 2) utilizing 0.3-mm-thick electrodes. The thermocouples, spaced 100 mm apart, were embedded ftush with the porous tube wall. Bach thermocouple junction was located on the internal tube surface. During the experiment, measurements were obtained for the porous wall temperature, screen temperature, inlet and outlet ftow temperatures, ftow rates of nitrogen and helium along the tube and for that drawn through the wall into the peripheral channel, current intensity, and voltage to the heaters of the porous wall and screen. The internal porous wall and screen temperatures were kept constant by regulating the voltage supplied to the test tube and screen. Thus, there was no radial temperature drop in the space around the porous tube. This arrangement permitted easy estimation of the heat ftux. The power input to the porous tube was removed partially by the helium ftow along the tube axis and partially by the fluid drawn through the wall; the latter is given by q = MsCp a r. where M. is the mass of the withdrawn gas, Cp is the heat capacity, and ar. is the measured temperature difference of the gas withdrawn before and after it enters the porous wall. t'L Tx T ' ~ 6 Fig. 3. Temperature distribution along porous tube for different rates of fluid withdrawal through the wall at Re; 0 = 2.4 x 104 • Legend: 1, G = 0.3 g/s; 2, G = 0.25 g/s; 3, G = 0.2 g/s; and 4, G = 0.15 g/s. 2 0 I I V / v V ~ V - _..... V 1 /~ ....., Ir:::: '=100 200 500 400 ~ I L.mm L. L. Vasillev, G. I. Bobrova, and L. A. Stasevieh 396 00 500 400 ~00 200 100 / 0.10 / V 0. 15 I I f/ 0.20 I I 0.25 C,g/s Fig. 4. Average Nusselt nurober along tube length versus rate of fluid withdrawal. When .1 T, is zero, as in this study, the heat ftux density was determined as the power per unit thermal surface of the porous tube. Figure 3 presents the experimental temperature distribution along the porous tube for different rates of fluid withdrawal through the porous wall when the helium ftow rate at the porous tube inlet was held constant at a Reynolds number of 2.4 x 104 • The Reynolds number in this case was computed in terms of the velocity at the porous tube inlet. The velocity was defined as V = V/ F, where V is the helium ftow rate at the inlet. The viscosity of helium and other physical parameters were evaluated at a temperature of 10 K. The inlet helium pressure was fixed at 1.1 atm while the outlet pressure was 1.05 atm. The Nusselt number was calculated in the following manner. The average heat transfer coefficient along the porous tubewas determined for each condition and the heliumthermal conductivity was estimated at 10 K. This average Nusselt number along the tube is presented in Fig. 4. With increasing suction through the wall, the temperature along the tube decreased. This was valid over the range of Reynolds numbers encountered (1.6 x 104 to 4.5 x 104 ). There was no marked effect of wall porosity on the temperature profile within the temperature range investigated. Temperature distributions along the tube and Nusselt numbers were compared with similar relations obtained by Aggarwal and Hollingsnorth [9 ]. The latter presented results for heat transfer in a porous tube with air and noted that at a fixed Reynolds number for the inlet ftow, the Nusselt number increases with increased suction. The temperature distribution along a porous tube and the Nusselt number as a function of the suction obtained by those authors are similar to the data presented in Figs. 3 and 4. In other words, the suction effect on the temperature distribution and the heat transfer coefficients for air and helium appear to be the same. NOTATION G= I= M, = Nu = quantity of heliuro drawn through porous wall porous tube length roass of gas withdrawn Nusseil nurober q = heat ftux density Re = Reynolds nurober T = teroperature of internal porous tube surface T0 = starting length teroperature V = heliuro ftow rate at porous tube inlet F = cross-sectional area of porous tube Heat Transfer of Helium in a Pipe with Suction 397 REFERENCES J. R. Sellars, J. App/. Phys. 26:489 (1955). A. S. Berrnan, J. Appl. Phys. 24:9 (1953). S. W. Yuan and A. B. Finkelstein, Trans. ASME 78:4 (1956). R. Kinney, Intern. J. Heat Mass Transfer 11:9 (1969). R. Y. Pederson and R. Kinney, Intern. J. Heat Mass Transfer 14:1 (1971). G. Baithby, Intern. J. Heat Mass Transfer 14:1 (1971). R. Kinney and E. Sparrow, Trans. ASME (Ser. C) (2):121 (1970). G. I. Bobrova and V. A. Morgun, in Heat and Mass Transfer af Cryagenic Fluids in Paraus Heat Exchangers, Izd. ITMO AN BSSR, Minsk, U.S.S.R. (1974), p. 63. 9. J. K. Aggarwal and M. A. Hollingsnorth, Intern. J. Heat Mass Transfer 16:591 (1973). 1. 2. 3. 4. 5. 6. 7. 8. G-7 BEAT TRANSFER AND HELIUM REPLENISHMENT IN CABLED CONDUCTOR COOLING CHANNELS* P. F. Michaelson, R. Quay, and R. F. Koenig General Electric Company Schenectady, New York and P. L. Walstrom and J. S. Goddard Oak Ridge National Laboratory Oak Ridge, Tennessee In the design of the superconducting conductor for the Large Coil Program (LCP) test coil, which is a large magnet cooled by pool boiling in liquid helium and operated with its bore axis horizontal, it is necessary to know the heat transfer capabilities of the conductor as a function of orientation (i.e., location along the coil perimeter) in the hydrodynamic environment of the magnet winding. In addition to conductor orientation, the conductor configuration and winding ventilation might be expected to affect the rate of removal of Joule heat from the stabilizer. This rate is important in the event of an incident which drives the conductor normal because the rate must be high enough to permit the conductor to recover to its superconducting condition. Early research by others [1-6] suggested that allowable heat ftuxes for recovery from a normalcy in a horizontally oriented conductor surface region of the magnet winding may be half or less of those of vertically oriented regions. This test was pedormed to determine the effects of orientation on the steadystate recovery heat ftux, and on possible hysteresis in the helium boiling curve, for two sample configurations, each configuration consisting of a set or array of Large Coil Segment (LCS) test conductor bars, one set with and one set without holes through its cores. The LCS conductor is a prototype of the GE-LCP conductor, with cooling channels twice as wide as the LCP conductor's. The LCP conductor and the existing LCS conductor have stlbelements cabled and soldered about a copper core as shown in Fig. 1. The subelements are a composite of a superconductor element Iaminated into a grooved copper stabilizer. The LCS conductor was used for this test because it is similar to the LCP conductor and was available. * Work supported by the U. S. Department of Energy under contract No. W-7405-eng-26 with the Union Carbide Corporation. 398 Heat Transfer and Helium Replenishment in Cabled Conductor Cooling Channels 399 Fig. 1. LCS conductor. TEST PROCEDURE The test sample was a three column by ten row array of 30.5-cm-long dummy (nonsuperconducting) LCS conductor bars with insulating spacers between rows and columns, simulating the GE-LCP test coil winding cross section (Fig. 2). The sample is long compared to the important cooling channellength (13 cm); this is the distance between intersections of the subelement grooves and the grooves in the insulating spacers between columns (Fig. 3). The sample was prepared at GE, and sent to ORNL where it was instrumented and tested. Heaters were installed in two of the conductor bars. The heaters, magnesium oxide insulated Niebrome wires with an outer sheath of Inconel, were soldered into the fulllength of the groove in ten of the eleven subelements of each of the two heated bars. In one of the two heated conductor bars, transverse holes approximately 2 mm in diameter were drilled through the copper core between subelements to enhance helium circulation. The turn-to-turn insulation and the LCS bars above and below this conductor were also provided with transverse holes with a random spatial relationship to the holes in the heated conductor bar. Differential Au 0.07 at. %Fe vs. Cu thermocouples were then installed. The reference junctions were located in the helium bath below the test assembly, to avoid vapor from the heated bars. A total of twelve thermocouples were installed on each of the two heated bars; in addition, a thermocouple was installed on each of the four bars surrounding each of the two heated bars. The thermocouple locations on each of the heated bars included the heated subelements, the unheated subelement, and the core. After being instrumented, the bars were assembled into the steel support fixture, as shown in Fig. 2. The cover plate was bolted on using a torque such that the pressure on the test bars during the test would be the same as that expected on the conductors in the LCP 400 P. F. Micbaelson, R. Quay, R. F. Koenig, P. L. Walstrom, and J. S. Goddard t2 '" · HEATED SPECtMENS LCP PAN GAKE - TO-PANCAKE tNSULATION ~.._....-- GE SLOT OPPER CORE I HOLES ORILLED THROUGH CORE ( LOWER SPECIMEN ONL Y) Fig. 2. Sampie geometry, simulating the LCP test coil winding cross section. Sampie bar orientation shown is horizontal (angle= 0°). test coil durlog operation. This was done so that the heat transfer from the heated conductor to the adjacent unheated conductors in the test would simulate that in service in the test coil. The instrumented fixture was mounted on a pivot mechanism attached to the dewar header (top ftange), which allowed the sample orientation tobe changed to known angles from outside the dewar. The apparatus was precooled with liquid nitrogen, then cooled with helium. Data of the heat ftux vs. temperature difference between the various sample locations and the helium bath were measured by applying power to the heaters while monitorlog the thermocouple outputs. Power data and thermocouple voltages were recorded by a fast multichannel digital system after amplification with differential amplifiers (Fig. 4). The heater current was programmed in either a square-wave shape or a staircase-up, staircase-down waveform (Fig. 5). Each step in the heater power staircase was three seconds in width. This duration was sufficiently long that steady-state temperatures were reached by all thermocouples, with rare exceptions, before a change to a new heater power level (Fig. 6). The measurements were taken at both increasing and decreasing heater power steps in order to detect any hysteresis in the liquid helium boiling curve. The steady-state helium boiling curve for increasing and decreasing heater powerlevelswas measured for one sample set at one orientation by this process. The Heat Transfer and Helium Replenishment in Cabled Conductor Coollog Cbannels Fig. 3. Majordetails oftheLarge Coil Programtest coil. 401 P. F. Mlchaelson, R. Quay, R. F. Koenig, P. L. Walstrom, and J. S. Goddard 401 16 THERMOCOUPLE 4 CHANNEL BIOMATION TRANSIE NT RECORDER X-Y RECORDER CHANNELS .... 0 < ...J .., ~ Cu >----"'+l ~ u Au-Fe TEST SPECIMEN POP 11/03 MICROCOMPUTER FLOPPY DISC <t cD CURRENTSENSING RESISTOR MICROPROCESSORFUNCTION GENERATOR Fig. 4. Schematic of measurement apparatus. Fig. 4. Schematic of measurement apparatus. Fig. 5. Example of heater power Ievel vs. time ; power Ievel at maximum step is 145 W for this run. Heat Transfer and Helium Replenishment in Cabled Conductor Cooling Channels 403 Fig. 6. Example of thermocouple voltage vs. time; l:lT at maximum step is 8.5 K for this run. process was repeated for both sample sets, with each heated separately, for each angle of orientation. The recovery heat ftux was calculated by dividing the heater power at the point of abrupt slope change ("knee") in the boiling curve by the product of the length and cooled perimeter of the heated bar. The LCS conductor has a cooled perimeter of 100 mm, located along the wetted surfaces of the cooling channels. These channels are 3 mm wide by 3 mm high by 130 mm long. The entire boiling curve is required for any complete stability analysis of a conductor, but the recovery heat ftux may be used for a quick comparison between experimental results. TEST RESULTS AND DISCUSSION The steady-state helium boiiing curves obtained in this test are presented in Figs. 7 and 8. Also given in each figure are the corresponding values for the recovery heat ftux vs. angle of orientation of the conductors. Very few of the heat steps displayed any hysteresis and the amount was small (a few percent or less) when present at all. "Hysteresis" herein is defined as different heat transfer rates at a given tl T between the sample surface and the coolant bath, depending on whether the sample temperature was rising or falling with time. Temperature gradients along the heated sample as weil as across the solder interfaces between the subelements and the core were small (a few percent or so) compared to the temperature difference between the sample and the helium bath, even though the ends of the heated bar were not insulated from the helium bath. At high power Ievels in the horizontal orientation, there was some heating (0.1 to 0.3 K) of the bars immediately above the heated bar. The magnitude of the experimental uncertainty in heater power vs. temperature is approximated by the size of the open circles used for each data point shown in Figs. 7 and 8. The uncertainty in the heat ftux values is estimated as ± 10%, with 404 P. F. Mlthaelson, R. Quay, R. F. Koenlg, P. L. Walstrom, and J. S. Goddard 180 SAMPLE WITH NO TRANSVERSE HOLES 160 9 2 • Omm (W/cm ) 0 . 18 0 . 22 0 .24 0 . 25 140 ~ 5 ~ 120 100 0 0. er: 80 w f- <I w 60 I 40 20 0 2 0 3 7 6 5 4 9 8 1t tO 6T , K Fig. 7. Steady-state helium boiling curve; sample without transverse holes. approximately half of this arising from the uncertainty in the cooled perimeter of the LCS conductor, and half arising from uncertainty in the amount of heat conducted to the adjacent bars and off the ends of the bars. A study of transient heat transfer effects may help in assessing the likelihood of formation of a temporary normalcy. Such data are useful in general, but arenot needed for the LCP test coil, so this test was not designed to systematically study transient effects. However, some general or qualitative information may be extracted 180 SAMPLE WITH TRANSVERSE HOLES 160 140 ~ t20 cr: ~ tOO 0 0. er: w 80 I- <I w • 9 60 Om;n ( W/cm o· I 10• 0 .27 20 20° 45" 0 .30 0 . 33 0.33 90" 0 2 4 5 6 7 ) 0.26 40 0 2 8 9 10 II 6T,K Fig. 8. Steady-state helium boiling curve; sample with transverse holes. Beat Trausfer and HeUum Replellishment in Cabled Conduetor CooUng Chaneis 405 from the data. For example, most of the heater Ievel steps produced thermocouple signals with no visible transient content, and this is true of the example signal of Fig. 6. However, a few of the square wave traces, taken with a high data sampling rate, did show transients for both of the samples (holes and no holes in the core), and for time durations that were "hydrodynamic-like," i.e., large fractions of a second. Transitions from nucleate to film boiling heat transfer rates, on the order of 0.1 s, were also observed. In interpreting this (transient case) information, it is noted that mechanisms involved in the transient heat transfer may include as parameters: (1) the enthalpy of the copper and/or the helium, (2) the latent heat of the helium, (3) the establishment of a heat-induced ftow pattern in the helium channels, and (4) possible long thermal time constants across solder interfaces of low thermal conductivity. Mechanisms (1) and (2) appear unlikely, based on the long time duration of the transients observed. Mechanism (4) appears unlikely, based upon the small thermal gradients measured during steady-state heating. The data are suggestive of a velocity-enhanced heat transferrate which occurs after a helium ftow pattern is set up by vapor jetting from the heated region. Further study of transient heat transfer in conductor bondies appears warranted. CONCLUSIONS Changing the orientation of a simulated LCP winding section (a sample of "complex" geometry, with surfaces oriented in many directions) from vertical to horizontal degrades the recovery heat ftux from 0.28 W /cm 2 for vertically oriented conductor bondies to 0.18 W /cm 2 for horizontally oriented bundles. Also, the addition of transverse holes in the insulation and conductor core enhances the heat transfer. Based in part on this data and results from a simple, long-channel-type heat transfer experiment a less conservative LCP design was adopted. (Previously an 80% degradationbad been assumed.) In another "complex-sample" experiment [8 ], an angle-dependent degradation of 33% was reported for another design of a conductor for the Large Coil Program. The higher transient and steady-state heat transfer rates obtained using complex samples, as compared to simple long channel or ftat plate data, may be due to coolant ftow that is induced by the heating, as has been suggested for other restricted bath geometry conductors [9 ' 10]. n, REFERENCES 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. B. J. Maddock, G. B. James, and W. T. Norris, Cryogenics 9:261 (1969). A. P. Butler, G. B. James, B. J. Maddock, and W. T. Norris, Intern. J. Heat Mass Transfer 13:105 D. N. Lyon, in Advances in Cryogenic Engineering, Vol. 10, Plenum Press, New York (1965), p. 371. G. B. James, K. G. Lewis, and B. J. Maddock, Cryogenics 10:480 (1970). M. N. Wilson, in Liquid Helium Technology, Pergarnon Press, Oxford, England (1966), p. 109. D. N. Cornish, D. W. Deis, A. R. Harvey, D. G. Hirzel, J. E. Johnston, R. L. Leber, R. L. Nelson, and J. P. Zbasnik, in Proc. 6th Intern. Conf on Magnet Technology, ALFA, Bratislava, Czechoslovakia (1977), p. 76. F. J. Reles, J. P. Heinrich, and R. E. Schwall, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 406. R. P. Krause, E. H. Christensen, R. D. Bradshaw, and R. E. Tatro, IEEE Trans. Magn. Mag-15:748 (1979). J. W. Lue, J. R. Miller, and L. Dresner, in Advances in Cryogenic Engineering, Vol. 25, Plenum Press, New York (1980), p. 251. M. 0. Hoenig, A. G. Montgomery, and S. J. Waldman, IEEE Trans. Magn. Mag-15:793 (1979). MEASUREMENTS OF BEAT TRANSFER AND HELIUM REPLENISHMENT IN LONG NARROW CHANNELS R. E. Schwall, F. J. Reles,* and J. P. Heinriebt Intermagnetics General Corporation Guilderland, New York INTRODUCTION The use of large cabled conductors has proven to be a cost effective method of fabricating high-current, superconducting composites with low normal state heat flux. Such conductors are presently being manufactured for use on very large superconducting coils. In the course of work on the design of severallarge coils, a need was identified to expand the existing data base on boiling helium heat transfer in long narrow channels. The work reported here was conducted to obtain the data necessary for efficient conductor design. Of the small amount of data available in the literature, perhaps the best-known which presents critical heat ftux data work isthat of James, Lewis, and Maddock for channels up to 2 cm in length. Wilson [ ] measured channels up to 20 cm in length, but the heated surface in his experimentwas stainless steel and there is some question about the applicability of the results to copper channels. Sydoriak [3 ] presented a hydrodynamic correlation of data on 72 different parallel plate evaporators with the plates vertical in all cases. In the present experiment, the heat flux as a function of channel orientation has been measured for a single channel size (13 x 0.11 x {).37 cm) which is typical of those utilized in large cabled conductors. Partiewar attention was given to the variation of the boiling characteristic for orientations near horizontal and to the reversibility of the heat transfer curve. The effect of helium replenishment was qualitatively evaluated by obtaining data with manifolds to restriet the flow of helium into and out of the channels. CJ. EXPERIMENTAL APPARATUS AND PROCEDURE The experimental apparatus (see Fig. 1) consisted of a copper plate with grooves milled on one face. Ten grooves were cut with a center-to-center spacing of 0.49 cm. The plate was mounted in a housing constructed of G-10 epoxy-fiberglass plates and • Present address: Xciton Corporation, Latham, New York. t Present address: General Elec:tric Company, Schenectady, New York. 406 Heat Transfer and Helium Replenisbment in Long Narrow Cbannels 407 all surfaces of the G-10 which contacted the copper plate were sealed with silicone rubber. G-10 thickness was 1.59 cm on the ungrooved face and the edges and 0.635 cm on the grooved face. The helium manifolds have a totallength of 3.56 cm and are angled at 45° such that those above the plate always opened upwards and those below the plate always opened downwards. Heater strips were mounted on the ungrooved face of the plate with current supplied via NbTi superconductive wire. This arrangement eliminated vapor generation near the sample due to heating of the Ieads. The heaters were powered by an IGC model 180M power supply with adjustable ramp rate. The power was determined by measuring the current and the valtage across the heater. Temperature was measured by a combination of resistive thermometers and Au-0.07 at.% Fe vs. Chrome I thermocouples. Two resistance thermometers were mounted in holes drilled in the sides of the plate perpendicular to the grooves. The sensors were located approximately 0.13 cm below the bottarn of the grooves and were mounted with heat sink compound to ensure good thermal contact. Thermocouples were installed on each corner of the plate and at the center top of the center tooth. In each case, the thermocouple junction was mounted on the copper plate and covered with silicone rubber. The thermocouple and thermometer Ieads were heat sinked to the side of the plate using cigarette paper and 7031 varnish. The thermocouple reference junctions were placed on a copper block located weil below the rest of the apparatus. Allleads were coated with silicone ruhher to reduce heat transfer to the bath. The angular orientation of the apparatus was varied by rotation about a simple COPPER BLOCK MANIFOLD ·r;:: - -+--- MANIFOLD 1 2.29 cm X 0 .48 cm l HEATER STRIPS Fig. 1. Heat transfer apparatus. Configuration illustrated is Case II channels-up. G10 HOUSING 408 R. E. ScbwaU, F. J. Reles, and J. P. Heinrich hinge at the base of the G-1 0 sheet. A threaded rod attached to the front of the G-1 0 housing and extending to the top of the dewar was used to lock the apparatus at any given angle. With this arrangement, the sample orientation could be varied from horizontal (6 = 0) to vertical (6 = 90°) without being removed from the dewar. To change the orientation from "channels-up" to "channels-down" however, required removal from the dewar. The experimental procedure consisted of slowly ramping the input power to some Ievel and then holding it constant to establish equilibrium. The readings of all the temperature sensors were recorded and the process was repeated until a temperature was reached that was weil into the film boiling regime. The power Ievel was then reduced in steps, and data points were again recorded. One of the thermocouple outputs was amplified and fed to an X- Y recorder along with the heater current so that the behavior of the sample could be continuously observed. EXPERIMENTAL RESULTS Heat transfer data were taken with the copper plate vertical, horizontal, and at a number of angles between these extremes. Particular emphasis was placed on characterizing the heat transfer at small angles to the horizontal, where the behavior was expected to change rapidly. Both orientations of the plate (channels-up and -down) were studied. All data were obtained in free-boiling helium at 4.2 K and 1 atnt pressure. Data were obtained for three experimental configurations: Case I: The grooved ends of the copper plate were exposed to the helium bath. Case 11: The grooved ends were covered with a 0.16-cm-thick G-10 plate with inlet and outlet manifolds in place. These manifolds, shown in Fig. 1, are intended to simulate the passages which would be available for vapor removal and helium replenishment in a large magnet. Case III : All channels and holes were blocked with silicone rubber to obtain a measurement of the background heat leak. A typical curve of heat flux vs. temperature difference, corrected for background heat loss, is shown in Fig. 2. Individual data points within a run were repeatable to ±2%, although, as discussed later, variations from run to run were somewhat larger. Initially, the heat flux increased rapidly at small temperature differences in the nucleate boiling regime. At a heat flux Ievel characteristic of the sample orientation, film boiling began and the surface temperature rose by several degrees. As the heat flux was increased further, the surface temperature increased approximately linearly with heat flux. When the heat flux was decreased, the same curve was retraced. This Iack of hysteresis, which is markedly different from typical 0 .4 i ~ 0 .3 ;t .rj 0.2 0 .1 0 .0-1'----+--+---+--+-----t--+---i-8 0 6 2 4 b.T , K Fig. 2. Heat flux vs. temperature difference for Case I, open ends. Channels-up 8 = 90•. Heat flux is corrected for background Iosses through G-10. Heat Transfer and Helium Replenishment in Long Narrow Channels 409 0 .5 45° 0.4 I E u J: 0 .3 ·Ö 0.1 0 . 0 +-----~--~r---~----~-----+-----+-----r- 0 2 4 8 6 10 12 14 AT, K Fig. 3. Heat flux vs. temperature difference for Case II channels-up. Parameter is orientation angle of channels. pool boiling, is apparently characteristic of long, narrow channels and/ or complex channel geometries. lt is qualitatively similar to that reported by Wilson [2 ] for channels of similar lengths and widths and by Cornish et al. [4 ] for an assembly of MFfF conductors from the MFTF magnet. The shape of the curves, even at the low-temperature end, is virtually identical to that reported by Wilson. In Fig. 3 the heat flux data are presented as a function of angle for Case II in the channels-up orientation. The effect of increasing angle is seentobe an increase in the breakaway heat flux and an increase in the heat flux attained for any given !1 T after breakaway is achieved. Figure 4 presents the same data for the channels-down orientation. The results are qualitatively similar with the exception of the fJ = 90° 0.5 30° 0 .4 jE 0 .3 u J: ·Ö 0.2 0 .1 0.0 0 2 4 6 A T, K 8 10 12 14 Fig. 4. Heat flux vs. temperature difference for Case II channels-down. Parameter is orientation angle of channels. 410 R. E. Schwall, F. J. Reles, and J. P. Heinrich 0.3 "I E u ~ ·O o Channels Up • Channels Down 0 .1 0 . 0 +:-----.;fc::-----::t:-:---+-,---+--+---:t:-:00 15" 30• 45" so• 75• 90" Hori zonta l Orientation of Channels Fig. 5. Breakaway heat flux as a function of channel orientation for Case I, open ends. Heat flux is corrected for background Iosses through G-10. curve, which is depressed below the () = 30° and () = 45° results. Individual data points have been omitted for clarity and the curves shown represent the best fit to a minimum of eight data points. Figures 5 and 6 show the breakaway heat ftux as a function of angle for two different experimental conditions. In each case, data for channels-up and channelsdown orientation are presented. The curves are qualitatively similar and resemble the data of James, Lewis, and Maddock CJ for short channels. There is a clear minimum at () = 0, i.e., forahorizontal channel and a fairly smooth variation of heat ftux with angle up to 90°. The minimum is most pronounced for Case II (Fig. 6) where the manifolds were placed over the helium inlet and outletslots of the copper plate. BACKGROUND LOSSES The data presented have been corrected for the "background" heat loss by conduction through the G-1 0 blocks, Ieads, and thermocouples. This contribution was measured at the conclusion of the experiment by filling all helium passages with silicone rubber (Case 111) and measuring in the channels-up configuration. The background loss exhibits an angular dependence, being a minimum at () = 0 and a maximum at () = 90°. At the breakaway transition, the background corresponds to approximately 6 x 10- 3 W /cm2 at oo and araund 2.7 x 10- 2 W /cm2 at 90°. These represent quite small corrections in an experiment of this type. DISCUSSION As noted in Figs. 5 and 6, the shape of the heat ftux vs. angle curves is similar for faces-up and faces-down. There is, however, a difference of up to 12% in the heat I E u ~ ·0 • Channels Up • Channels Down 0 .1 o.oo• 15• Horizontal 30 • 45" so• 75" 90" Orientat ion of Channels Fig. 6. Breakaway heat flux as a function of channel orientation, Case li. Helium manifolds in place. Heat flux is corrected for background Iosses through G-10. Heat Tnnsfer and Helium RepleDisbment in Long Narrow Channels 411 ftux measured. This is true even at 90°, where the cases should be identical. This discrepancy, although not large for a helium heat transfer measurement, is considerably greater than the precision of the measurements and its source is not understood. Also surprising is the fact that the breakaway heat ftux is generally larger in the channels-down configuration than in the channels-up mode. The work of Lyon [5 ] on ftat plates, for example, showed peak nucleate boiling heat ftux in an upward-facing ftat plate to be four times that in a downward-facing plate. The asymmetry in the thickness of the G-10 plates in the present apparatus could mean that the background losses would be smaller in the channels-down position than in the measured channels-up orientation. This would serve, however, to raise the corrected channels-down data and does not explain the present results. It is possible, however, that Lyon's data are not applicable in the present case. James et al. CJ reference the work of Lyon but their data show breakaway heat ftuxes that are higher in some cases with the heated foil facing downward. CONCLUSIONS The heat transfer to helium in narrow channels similar to those in large cabled conductors lacks hysteresis and the breakaway heat ftux is below the peak typically reported for pool boiling. The breakaway heat ftux and the heat ftux in the film-boiling region is found to depend strongly on the angle between the channel and the horizontal. The heat ftux obtained at any given angle is decreased and the heat ftux minimum at () = 0 is accentuated when the entry and exit of fluid to the channels is controlled by external manifolds. ACKNOWLEDGMENTS This work, performed at Intermagnetics General Corporation, was sponsored by the General Electric Company, Schenectady, New York, under ESPD Subcontract No. F91-51164. The authors gratefully acknowledge a nurober of helpful comments by one of the referees. REFERENCES 1. G. B. James, K. G. Lewis, and B. J. Maddock, Cryogenics 10:480 (1970). 2. M. N. Wilson, in Liquid Helium Technology, Pergarnon Press, Oxford, England (1966), p. 109. 3. S. G. Sydoriak, in Proc. 13th Intern. Conf. on Low Temperature Physics-LT-13, Vol. 4, K. D. Timmerhaus, W. J. O'Sullivan, and E. F. Hammel, eds., Plenum Press, New York (1974), p. 607. 4. D. N. Cornish, D. W. Deis, A. R. Harvey, D. G. Hirzel, J. E. Johnston, R. L. Leber, R. L. Nelson, and J. P. Zbasnik, in Proc. 6th Intern. Conf. on Magnet Technology, Editorial Committee MT6, ALFA Bratislava Publishers, Bratislava, Czechoslovakia (1978), p. 76. 5. D. N. Lyon, in Advances in Cryogenic Engineering, Vol. 10, Plenum Press, New York (1965), p. 371. G-9 VAPOR LOCKING AND HEAT TRANSFER UNDER TRANSIENT AND STEADY-STATE CONDIDONS* C.-J. Chen, S.-T. Wang, and J. W. Dawson Argonne National Labaratory Argonne, fllinois INTRODUCTION The technology of stable superconducting magnets has become synonymous with the study and use of composite conductors. The composite conductor, a superconductor paralleled with a normal metal, helps provide mag11et stability by supplying alternate electrical and thermal paths for the superconductor when it becomes normal. If these alternate paths of normal metal can carry the total transport current continuously and still remain below the transition temperature of the superconductor, the composite conductor is said tobe cryostable. The operational definition of cryostability requires sufficient cooling to dissipate Joule heating. The degree of cryogenic stability depends on the heat transfer characteristics of the liquid helium cooling channels. Normal zones created following mechanical disturbances will either grow or collapse depending on the heat transfer rates from the conductor to the adjacent cooling channels. It is important to design large magnets with cooling channels sufficiently large so that vapor binding will not occur under both steady-state and transient conditions. Steady-state and transient heat transfer to liquid helium channels has been studied by various investigators [1- 5 ]. This study was undertaken to determine the vapor locking in the cooling channels of a cryostable superconducting magnet rJ, to investigate the heat transfer characteristics under steady-state and transient conditions, and to study the effects of vapor accumulation of the multiple coillayers. Heat ftux, energy density, temperature rise, and vapor fraction, are defined in the following figures and paragraphs as electrical power to heaterI cooling surface area of the conductor, electrical energy to heaterI conductor volume, temperature difference between the sample and the liquid helium bath, and volume fraction of vapor in the cooling channel, respectively. APPARATUS AND MEASURING TECHNIQUES To obtain the effects of transient and steady-state heat transfer and vapor formation on cryostable conductors, samples were fabricated to simulate the real • Work supported by the U. S. Department of Energy. 41Z Vapor Locking and Heat Transfer under Transient and Steady-State Conditions 413 cryostable superconductor and the cooling channels to be used in the large MHD superconducting magnet, designated as CFFF-SCMS [6 ], currently under construction at Argonne National Labaratory (ANL). The cross section of the conductor is 3.1 by 0.47 cm and that of the cooling channel is 0.97 by 0.076 cm. Figure 1 shows an assembly of a simulated single coillayer. Three conductors are sandwiched in between insulation of 0.064-cm-thick pultruded fiberglass strip. A 0.0064-cm-thick double-coated adhesive Mylar tapewas used to bond the conductor together. The resulting assembly provided actual cooling channels. The assembly was covered about 50% with 0.64-cm-thick Micarta strips to simulate identical cooling conditions of the conductors within the CFFF -SCMS coil structures. A 0.0025-cm-thick stainless steel heater was inserted in the middle conductor. The insulation between the heater and the conductor was a 0.0075-cm-thick lens paper impregnated with GE-7031 insulating varnish. Maximum power dissipated by the heater was about 5 kW. The temperature of the conductors was measured with a Chromel vs. gold-0.07 at.% iron thermocouple. Good thermal coupling of the thermocouple insulation was provided by wrapping the indium solder tip of the thermocouple with a single layer of 0.0025-cm-thick lens paper impregnated with GE-7031 varnish. Fig. 1. Assembly of simulated single coillayer for vapor-locking and heat transfer experiments. Legend: A, simulated conductor with heater; B, simulated conductor without heater; C, pultruded fiberglass insulation and front-cooling channel; D, pultruded fiberglass insulation and rear-cooling channel; E, Micarta blocks; F, half-covering Micarta blocks; G, copper holders; H, horiznntAI cooling channel; and I, vertical cooling channel. 414 c..J. Chen, S.-T. Wang, and J. W. Dawson The vapor fraction in the channel was determined by measuring the change in the capacitance of the channel. The capacitance change, aC, of the channel due to the presence of vapor is given by (1) where a is the volume fraction of helium vapor inside the channel and CL is the capacitance of the channel when it is filled completely with liquid helium. When the channel is completely filled with helium gas, the vapor fraction, a, is equal to unity and the capacitance change reaches a maximum value designated as aCmax· If the temperature of the helium vapor remains constant, (1) can be simplified to ac a =-- aCmax (2) A capacitance bridge with a triaxial cable to compensate for any leakage current was used to measure the capacitance change. The sensitivity of this bridge was about 3 V /pf capacitance change. The maximum capacitance change for the cooling channel of the simulated assembly was about 1.4 pf. The wetted area of the front side of the middle conductor was 33.98 cm2 and that of the rear side was 16.16 cm 2 • The volume of the conductor used was 28.53 cm 3 while the volume of each channel was 2.58 cm 3 • RESULTS The experimental study was conducted employing either a steady state or a transient current to the heater. In either case, the temperature difference and the capacitance changes were measured. The capacitance bridge was calibrated before each run. Sufficient equilibration time was allowed after each energy pulse to ensure escape of all vapor bubbles from the channels and cooldown of the conductors to liquid helium temperature. The waiting duration is on the order of a minute and is proportional to the injected energy. Steady-State Results The steady-state heat transfer characteristics for the channel are shown in Fig. 2. This figure indicates that the critical heat ftux for the transition from nucleate boiling to film boiling is about 0.4 W I cm 2 while the recovery heat ftux for the transition from film boiling to nucleate boiling is about 0.25 W /cm 2 • The temperature rise changes quite markedly from 0.2 K to about 7 K when the transition from nucleate boiling to film boiling occurs. During the recovery process, the temperature decreases rapidly from 1 to 0.1 K when the nucleate boiling regime is attained. This indicates that a single-layer assembly can remove steady-state heat ftuxes up to 0.4 W /cm 2 , which is equivalent to the conductor dissipating a steady-state Joule heating of 1.16 W /cm length of conductor. The steady-state heat transfer coefficient, h, decreases shortly after the peak nucleate heat ftux is reached. This results in a decreased heat ftux to the channels. Figure 3 shows the vapor fraction observed at the front and the rear channels as a function of the applied heat ftux. The vapor fraction is about 0.44 for the front channel and about 0.36 for the rear channel when the heat ftux is close to the peak nucleate heat ftux. This is equivalent to about 0.1 cm 3 of liquid helium vaporized within the channels per centimeter length of conductor. _10_, 0 10 _, 10° TE MPERATUR[ RISE. 6 T. K - -- .,.._ 0 10' .; -_::_:;? Fig. 2. Heat transfer characteristics under steady-state conditions. :r: w- <(0 o-• "- ::> -' X ~ -. u w w ,_ ci uo -.....3': ":::;0 ~ r-----------------------------------. 0 ~ "'c:i 0 ~ c:i "' l>:;; _ c:i z HEAT FLUX INJECTED, W/CM' ' ""'10_, , , """'10-· ' , """'10° I Fig. 3. Characteristics of vapor formation under steadysiate conditions. = Ul ~ .... i" = "' SI ~ ~ i ~I10 _, .<, 0 0 '1 i :. ~· a ~ ~ Cl. = :::1 ~ :::1 =. ~ ä= Cl. , / :. IIQ :::1 9: 0 t"' ~ i c:i 0 ~ ~ c:i ~:; o., "' o. .... Q;o ..._ <( "l ug ;::: f RONT - ~RE A R ~ ,-----------------------------------------------~ 0 416 C.-J. Chen, S.-T. Wang, and J. W. Dawson Tnmsient Results For the transient heat transfer sturlies the pulse duration and power Ievel of the heater were varied as desired. The vapor fraction and temperature rise presented are the maximum values observed for each pulse. The heat transfer characteristics are presented below. Figure 4 shows the energy density vs. temperature rise for different pulse durations. These results indicate that the temperature rise is a function of the energy density and is independent of the pulse duration. The critical energy density for boiling transition is about 100 mJ/cm 3 • If the mechanical disturbance dissipates an amount of energy less than the critical value, the peak temperature rise is less than 1 K. When the injected energy density was increased to about 1 J I cm3 , the temperature of the conductor rose to about 15 K peak value. The temperature recoveryrate was about 0.02 K/ms. Figure 5 shows the heat ftux under different pulse duration vs. temperature rise. Observations showed that the shorter the pulse duration, the higher the critical value for boiling transition. When the pulse duration was Ionger than 200 ms, the heat transfer characteristics approached those observed under steady-state condition. These data show that the critical heat ftux for a 10-ms pulse is about ten timesthat of the critical heat ftux for a 200-ms pulse. The vapor fractions observed for the front and rear channels vs. the heat ftuxes under different pulse duration are plotted in Fig. 6. The vapor fraction of the front channel is less than that of the rear channel at the lower heat ftux range, but this trend reverses as the heat ftux is increased. However, the vapor fraction of the front channel again is less than that of the rear channel when the heat ftux has been increased sufficiently to induce the transition of nucleate boiling to film boiling. To explain these phenomena, consider the following arguments: 1. At the low-heat ftux range, both channels are in the free convection regime in which fluid moves under the inftuence of buoyant forces arising from changes in density. The velocity is zero at the heated surface (no-slip boundary condition), increases rapidly in a thin boundary layer adjacent to the surface, and becomes zero again far from the surface. The wetted surface of the front side continues from bottom to top, while the wetted surface of the rear side is blocked by turn-to-turn insulation. This obstruction inhibits the free convection in the rear channel. With these restrictions on convection, the vapor fraction in the rear channel would be greater than that in the front channel. 2. When the heat ftux increases, both channels are in the nucleate boiling regime. The vapor bubbles begin to appear at the heating surface and depart from the surface. Since the wetted area of the front side is about twice that of the rear side, the energy transferred into the front channel is more than that transferred into the rear channel. A larger transfer of energy results in a greater vapor fraction. 3. In the high heat ftux range, the liquid helium of the front channel apparently attains film boiling while the rear channel does not. Thus, the heat transfer into the rear channel is more than that into the front channel resulting in a vapor fraction that is greater for the rear channel than for the front channel. Figures 7 and 8 show the vapor fraction of the front channel and the rear channel vs. the energy density. The transition to film boiling occurred in the front channel in the 100 mJ/cm 3 range, but did notoccurin the rearchannel up to 1 J/cm 3 . Thisresult is consistent with the heat ftux data of Fig. 6. :::; 0 - " 10_, 10° TEMPERA TURE RISE , AT. K 10' + ~~~~~n;--,-,-rrnT~----~~~---r-r~ I DURAliON . - . DUR AliON ,.....,. DUR A!ION Fig_ 4. Heat transfer characteristics under transient conditions for different pulse durations_ _10_, 0 w~ zo w- 0:: <.;J >- 0 w z iii ~ ,_ E ....... -, u-- -c = 50 00 ms -c : 100.0 ms ot = 2000 ms IO.OOOms 1: : 0-o DURATION 2 ~------------------------------------, w g g 10I _, ö ! A; "C= ms 200.0 ms 1: = 100.0 TEMPERA TURE RISE, lH , K I """'I 10_, I " '""' 10 _, I ii illlll 10' I"""'' 10' I " "'"I 10 ' A =20.00 ms '( ::::: 50.00 I'I'IS 'C 1 ::::: 10.00 P'!'l l Fig_ 5_ Temperature rise of conductor with respect to heat ftux under transient conditions- :X: w <( .... LL. ...J :::> X ~- --,0 u wo I-- 0 ~ :::;; u g • ········• DURAl iON • --- - DURATION • - X DURA TION - D URATION cr---v OURATION 9r--------------------------------, ~ .... .... ..... a ~ t (""l i ~ lf Cl.. = ! [ ~ ..§~ ~ ~ ;: l 110 ~ ~ t"' ~ .. 'Cl 418 C.-J. Chen, S.-T. Wang, and J. W. Dawson 0 ~ ~--------------------------------------------, <>-<> f RON T 1 = 5.0 00 I"I"'S ~ ci $.000 .,..". REAR _ rRON T -:: = 10.000 ms x- ·x REAR 1 <>--<> fRON T (). -o RE AR t 't: ITI S = 10.000ms =100 .0 ms t= 100.0 T!'' :i/!1 I I I z I 0 ;:: r /X uo < ~ ci Cle ..... A/ 1 I I 6 I I ., I • I 0 ;(l ci 10° 10' HE AT FLU X IN JECTEO. W/ CM' Fig. 6. Vapor fraction of cooling channels with respect to heat flux under transient conditions. 0 .7 lOOm 0 .6 ,_"' ~ 0 ·:~ ~·9)r. 5 ... 0 .4 0.. ;! 0 . 3 0.2 0. 1 0 ( OO•• 20'" "'Y ~~.7 ~ 0 .5 ~ "' .... "'0 1'\ 501111 -~ .:Y I /l ID ·- [ f 100 1000 ENERGY DENSITY 10000 Fig. 7. Characteristics of vapor formation under transient conditions for the front channel. Legend : A, 1 ms; T , 2 ms; e. 5 ms ; 0 , 10 ms ; x, 20 ms; \1, 50 ms ; 0 , 100 ms ; • · 200 ms. Vapor Locking and Heat Transfer under Transient and Steady-State Conditions 0 .7 L ;t 0 .6 z 0 ~0 . 5 « a: ..._ a: 0.4 ~0. 0 .2 0 .I 0 ~ ....,7 ~ /"' V 10 / -- - - - ,j 100 419 - ' 1000 ENERGY OENSITY 10,000 OIJ / / tm' Fig. 8. Characteristics of vapor formation under transient conditions for the rear channel. Legend: &, 1 ms; ~ . 2 ms ; e, 5 ms ; 0 , 10 ms ; x, 20 ms ; 'V, 50 ms; 0, 100 ms ; • · 200 ms. CONCLUSIONS Based on the experimental results obtained, the following conclusions can be made: (1) The critical steady-state heat flux for the transition from nucleate boiling to film boiling is about 0.4 W I cm 2 ; (2) the critical transient energy density for the boiling transition is about 100 mJ/cm 3 ; (3) the front channel attains film boiling much easier than the rear channel; and (4) no temperature rise greater than 1 K is possible if the heat flux does not exceed the critical heat flux under steady-state conditions or the energy density does not exceed the critical energy density under transient conditions. REFERENCES 1. M. N. Wilson and Y. Iwasa, Cryogenics 18(1):17 (1978). 2. C. N. Whetstone and R. W. Boom, in Advances in Cryogenic Engineering, Vol. 13, Plenum Press, New York (1968), p. 68. 3. 0. Tsukamoto and S. Kobayashi, J. Appl. Phys. 46:1359 (1975). 4. G. B. James, K. G. Lewis, and B. J. Maddock, Cryogenics 10:480 (1970). 5. M. A. Hilal, J. W. Dawson, J. D. Gonczy, L. R. Turner, and S.-T. Wang, IEEE Trans. Magn . Mag-15:59 (1979). 6. S.-T. Wang, R. C. Niemann, L. R. Turner, L. Genens, W. Pelczarski, J. Gonczy, J. Hoffman, Y-C. Huang, N. Modjeski, and E. Kraft, IEEE Trans. Magn. Mag-15:302 (1979). G-10 FORCED TWO-PHASE HELIUM COOLING OF LARGE SUPERCONDUCTING MAGNETS M. A. Green, W. A. Borns, and J. D. Taylor Lawrence Berkeley Labaratory Berkeley, California INTRODUCTION A major problern with alllarge superconducting magnets is the cryogenic and refrigeration system. Almost all of the !arge magnets are cryogenically stabilized. They are cooled in a bath of boiling helium. Nucleate boiling in liquid helium permits heat ftuxes of 3000 W /m 2 to be transferred with a temperature drop of less than 0.5 K. The bath-cooled magnet is difficult to cool from room temperature to 4 K. Helium, which is a difficult fluid to use as a coolant because of its low atomic weight, must ftow into each region of the magnet; if the ftow in a section of the magnet is restricted, stratification occurs and that magnet section remains warm. Once the pool-boiled magnet is cold and the cryostat is filled with liquid helium, the refrigeration problemsarenot over. Helium has a very low heat of vaporization (20 JI g at 4.2 K) C]. As a result, the release of !arge quantities of magnet-stored energy during a quench (even cryostable magnets quench on occasion) will result in !arge quantities of helium gas being ftashed. For example, 1' MJ of energy will vaporize 400 liquid Iiters of helium into 286m 3 (10,000 te) of helium gas at STP conditions. The !arge amount of helium gas generated by a quench requires careful design of the inner cryostat vessel and the relief valve system. The thickness of a !arge diameter pressure vessel which is designed for at least 3 atm is not trivial. · Large superconducting magnets do not have to be bath cooled in many cases. lt is often immaterial how an intrinsically stable superconductor is cooled as long as it is kept below its critical temperature. Bath cooling, with liquid helium in direct contact with the superconductor, is undesirable in any magnet which is designed so that it can quench. The helium in direct contact with the superconductor restricts the growth of the propagating normal region, which makes burnout of the coil more likely. The helium cooling in an intrinsically stable coil should be away from the superconductor. A forced-ftow helium tubular cooling system will provide all of the cooling that is needed. This systemwill also avoid all of the major problems which are encountered in any !arge cryogenic system. The advantages of the tubular cooling over an ordinary bath-cooled system are as follows: 1. The cooldown of the magnet is weil controlled because the helium ftows in a well-defined path. 420 Foreed Two-Phue Helium Cooling of Large Sapereondaetlng Magnets 421 2. The mass of the tubular cooling system is less than that of a helium bath cryostat. 3. The amount of helium in direct contact with the magnet coil is minimized. Helium boil-off during a quench is orderly and weil controlled. Cryogenic safety in such a system is greatly enhanced. 4. Tubular cooling is unaffected by magnetic forces. A bath-cooled superconducting magnet in space is vulnerable to diamagnetic repulsion; a tubular cooled system is not. The tubular cooling system is not new; a number of superconducting magnets use this system 5]. The Lawrence Berkeley Labaratory system is different because it employs two-phase helium instead of supercritical helium (the critical pressure of helium is 2.25 x 105 Pa; the critical temperature for helium is 5.19 K). The reasons for choosing two-phase helium over supercritical helium are as follows: e- 1. A two-phase "boiling" helium system will operate at lower temperatures than a supercritical helium system. A lower Operating temperature allows the superconductor to transmit more current. 2. The mass ftow required in the cooling circuit for a given amount of refrigeration is lower for a two-phase system than for a supercritical system. In most cases the pressure drop will be lower as weil. 3. Boiling two-phase helium can transfer large local heat ftuxes without changing the temperature of the stream. Two-phase helium forced cooling is, in most cases, superior to supercritical cooling. There are instances when this is not true; they are when (1) the dielectric strength of the helium must be uniform (for example, in superconducting transmission lines); (2) the pressure drops in the system must be above 5 x 104 Pa; (3) the Operating temperature range includes temperatures above 5.2 K; and (4) when the superconductor is supposed to have full cryogenic stability. DESIGN OF A TWO-PHASE COOLING SYSTEM The most important consideration in design of two-phase ftow systems (this applies to other systems as weil) is the elimination of parallel paths. The most desirable system is the simple series ftow system. A series ftow system has almost no control problems; therefore, no additional refrigeration is expended to control the ftow. When a parallel ftow system cannot be avoided, one should strive to minimize the number of parallel ftow circuits. Each parallel circuit should have its own control dewar system. The principle of the control dewar will be discussed later. There are two other important considerations when one designs a two-phase helium ftow system. Theseare low pressure drop and ftow stability. The former is achieved by choosing an appropriate length-to-diameter ratio and the latter is achieved by choosing an appropriate operating regime. The two, unfortunately, go band in band. There is, however, wide latitude in choosing tube diameter and tube length. In general, one wants to operate on the liquid side of the two-phase liquid vapor dome. One also wants to operate in the bubble and froth regime of the Baker [6 ] diagram. (In general, one would like to operate the system and have mass ftows per unit area of greater than 10 kg/ s-m 2 , which avoids the slug-and-plug ftow regime [1].) Two-Phase Pressure Drop A series of pressure drop calculations were performed using the MartinelliNelson Technique [8 ]. Although the calculations agreed with the experimental 422 M. A. Green, W. A. 81111111, ud J. D. Taylor results, they were very time consuming. A much simpler equation, given below, yielded almost the same results for round tubes (the reason for this isthat the density change across the two-phase dome is less than a factor of 8): !l.P = _!_ m2 (1 + Tr2 P fz) _!_ D D4 (1) where f = 0.184 Re -o. 2 (turbulent ftow smooth pipes), Re = 4m/ TrD~J-, and where m is the mass ftow, p is the average density of the exiting helium, z is the tube length, D is the tube diameter, IL is the helium gas phase viscosity; and !l.P is the pressure drop along the tube. In order to achieve low pressure drop, one wants to maximize the density of helium in the cooling tube. For that reason, the systems are designed to operate so that liquid helium which circulates has the lowest quality X. (Quality, in this case, is defined in the samesense as it is for steam.) Quality Xis defined as (2) where H, S, and V are the enthalpy, entropy, and specific volume, respectively; the subscript I denotes saturated liquid, the subscript g denotes saturated vapor, while a quantity without a subscript denotes the two-phase region. H, S, and V are bounded by the liquid vapor dome so that the quality X is always between zero and unity (0 :S x :S 1). Once V, the two-phase specific volume is known, the density of the two-phase fluid can be determined by taking the reciprocal of V. Since a reduction in the quality for the ftow system reduces the pressure drop, the operating temperature is reduced and the temperature change across the system is minimized. Helium Control Dewar and Cool Dewar Circuit Two types of systems can be used to circulate low-quality helium through the magnet cooling tube, namely, (1) a liquidheliumpump used as a circulator or (2) a refrigerator compressor used as a circulator. Both systems use a heat exchanger to ensure that the helium will enter the system at or near the saturated liquid line. Figure 1 shows schematic diagrams for the two approaches. The heliumpump loop system shown in Fig. 1a has the following advantages: (1) The refrigerator is completely decoupled from the Ioad. In theory, one could Substitute liquid helium from a storage dewar for the refrigerator. (2) The mass ftow through the system is limited by the capacity of the pump, not the capacity of the refrigerator. Key disadvantages associated with the use of a heliumpump system are that the pump work is absorbed by the helium, requiring extra refrigeration, and the simple pump loop system cannot be used to cool the magnet from room temperature; in such a system one should connect the magnet cooling system directly to a refrigerator. The refrigerator compressor can be used as a circulator provided a heat exchanger is used with an accumulator. The function of this heat exchanger is to reduce the inlet quality to the magnet cooling tube. The quality change across the magnet remains constant for a given mass ftow in the circuit. The circuit shown in Fig. 1b has two Joule-Thomson valves (J-T valves) which expand the gas in two stages. The first J-T valve expands from 15-18 bar to 3 bar. The heat is transferred to the boiling liquid helium in the accumulator tank as the gas ftows through the heat exchanger. Expansion of the gas from 3 bar to the final inlet pressure to the magnet Forced Two-Phase Helium Cooling of Large Superconduding Magnets 423 REFRIGERATION MAGNET (a) LIQUID HELIUM CIRCULATION WITH PUMP (b) LIQUID HELIUM CIRCULATION WITH REFRIG. COMPRESSOR Fig. 1. Schematic of two types of two-phase helium circulation for tubular cooled superconducting magnets. coil will result in an inlet quality which approaches zero. If there were no heat exchanger, the inlet quality to the magnet coil would be around 0.4; the pressure drop in the tubular cooling system would be higher by a factor of 2 to 3. The system shown in Fig. 1b is analyzed from a thermodynamical standpoint by Green [9 ]. This system is capable of being operated at 40 to 60% over the capacity of the refrigerator for short periods of time. (The amount of time is dependent on the amount of excess liquid helium in the accumulator.) lt can be cooled by the refrigerator directly, provided one bypasses the gasback to the compressors after it has been circulated through the Ioad being cooled from room temperature. Figure 2 shows how the system shown in Fig. 1b can be inserted into a system to cool a magnet down, operate the magnet assuming a certain pressure drop, and warm the magnet back up to room temperature. Figure 2 is highly schematic, but it shows the essence of the LBL control dewar system which can be located remotely from the Ioad being cooled. A hybrid system which uses both a helium pump and refrigerator compressors as circulators has been built to cool the TPC magnet [1°]. Redundancy has been built into the system so that the magnet can be operated whether or not the refrigerator operates. (The helium pump circulates stored liquid while the refrigerator is inoperative.) If the helium pump is inoperable but there is a refrigerator, the forced flow through the coil is supplied by the refrigerator itself. I 300 K to 80 K. ~ Open J- T valve. 11KJ Closed full flow valve. Lines not corrying flow. Open full flow volve. Closed J- T valve. [:;o::J ....... Llnes carrying flow. CLEGEND> cl Normal operation at temperotures below 10 K. al Cooldown from ~--:;;~----+-4- Control Dewar * I iI r-- 1 Fig. 2. Cooldown, normal operation, and warmup of a superconducting magnet with forced two-phase helium cooling. dl Wormup from 4 K to 300 K. I...M/INVIIINM--- I \ Control I I bl Cooldown from 80 K to 10 K. l ~ ~ ~ [ j ~ ~ f ~ ~ e Forced Two-Phase HeUum CooUng of Luge Sapercondacting Magnets 425 Garden Hose Eflect Large superconducting magnets, that are force-cooled with two-phase helium, may face an additional problern which depends on the orientation of the cooling circuit. If the magnet cooling circuit is oriented so that the two-phase ftow is horizontal or near horizontal, the two-phase ftow should be stable and should result in a pressure drop that is predictable using (1). If the two-phase ftow is vertical with many up and down loops, the ftow is subject to the so-called "garden hose" pressure drop and "garden hose" oscillation. Such a pressure drop with associated oscillations is not ful!y understood, but it has been observed experimentally. The pressure drop can be visualized by looking at a garden hose which is half full of water and hangs in a coil on a peg. lt takes an increase in pressure to force the water out of the hose due to the additive effect of the various heads of water in each of the coils. A similar behavior has been noted in helium systems even though the ftow is presumed to be in the bubble-and-froth regime of the Baker diagram. This type of pressure drop can be estimated approximately by using the following expression for a coil of N round loops, of major diameter D, which have an axis orientation angle of () from the horizontal: M* = 0.2 Nda (p 1 - pg) cos 8 (3) where P* is the pressure drop due to the above effect, p, is the density of the helium liquid phase, pg is the density of the helium gas phase, and a is the acceleration of gravity. The coefficient 0.2 is derived from experimental measurements ofthistype of pressure drop and is probably related to the velocity of the coolant. At best, (3) can only be regarded as an approximation. M* is added to the pressure drop calculated in (1). From (3) it appears that this type of pressure dropwill decrease as the system approaches the critical pressure because the quantity (p,- pg) is reduced as one approaches the critical pressure. The other obvious Iimitation is that !l.P* is less than the difference between the critical pressure and the system exit pressure regardless of the values of N and D. A reduction of the ND product will reduce the pressure drop. Changing the axis of the hoops from horizontal (8 = 0) to vertical (8 = 90) will eliminate !l.P* entirely. This specific type of oscillation is probably caused by a change in ftow regime, i.e., from bubble and froth to slug and plug. Starting and stopping the ftow requires the ftow to cross over the slug-and-plug ftow regime. The oscillatory behavior of slug-and-plug ftow, coupled with various spring constants of the system, which includes the refrigerator or pump loop, probably contributes to the oscillation. The amplitude of the pressurepulse due tothistype of oscillation appears tobe no greater than the pressure drop M*. Gas-Cooled Electrical Leads Gas-cooled electricalleads present no problern on a forced-cooled two-phase helium. First, one can take two-phase helium directly from the cooling system without any ill effect to the magnet, the Ieads, or the cooling circuit. No additional Iead pot is necessary. Second, the superconducting Ieads from the magnet do not have to be in liquid helium. They only have to be in contact with the refrigeration supplied by liquid helium. 426 M. A. Green, W. A. Bums, and J. D. Taylor The implications of the two previous points permit one to design magnets that have the electricalleads brought out into the vacuum space and attached to a pipe or busbar cooled by liquid helium. The high-voltage arcing, which is a common fault in large magnets, can be eliminated. Therefore, quench protection voltages up to 5 kV are relatively easy to handle. A second implication is that installation and assembly can be simplified. The design of the electricallead itself is not changed by the fact that it operates with a two-phase cooling system. However, one must remernher that Iead operation on a refrigerator is different from operation in a bath of helium without a refrigerator. (This statement is true for bath-cooled magnets as weil as force-cooled magnets.) The precautions one must take with gas-cooled Ieads which operate with a forcedftow system are (I) the Ieads should be oriented so that the warm end is above the cold end, and (2) the ftow of gas through the Ieads must be controlled so that the Ieads do not operate at a warm temperature. One may use a controller which measures Iead temperature or one may measure the voltage drop across the electrical Iead. EXPERIMENTAL MEASUREMENTS The LBL group has operated three large test superconducting magnets using a two-phase tubular cooling system. Most of the early work was undertaken using a modified version of the circulation system shown in Fig. lb [ 11 ' 12 ]. Tests were performed in the summer of 1979 using a helium pump. Solenoid Tests on a Refrigerator Two 1-m-diameter test solenoids and one 2-m-diameter test solenoid were cooled by two-phase helium and were operated at or near critical current at temperatures from 4.6 to 4.8 K. The two I-m-solenoids were operated in series. The cooling circuit tubes are shown in Fig. 3. The circuit length was 235 m while the cooling tubeIDwas 10.8 mm. There were a total of 72 turns of cooling tube Im in diameter. The axis of the two coils in series was vertical (8 = 90 K). The 2-m test coil contains 365m of 10.8-mm-ID tube wound in 55 turns which are 2m in diameter. During the test the coil axis was vertical. The 2-m-diameter solenoid is shown in Fig. 4. The cooldown of both tests took a bit over one day when a CTi model 1400 refrigerator was used. Figure 5 shows the pressure drop in the cooling circuit as the two I-m-diameter solenoids cooled down. The onset of two-phase ftow in the tube was signaled by a sudden drop in the pressure drop across the circuit. Since the cooling circuit axis was vertical, no "garden hose" pressure drop or oscillation was observed. There was no oscillatory behavior observed in temperature or pressure. Changes in pressure drop were gradual. Increasing the liquid in the accumulator Fig. 3. Cross section of one of the LBL twophase cooled test coils. Forced Two-Phase Helium Cooling of Large Superconducting Magnets Fig. 4. Two-meter-diameter LBL test coil cooled with twophase helium. 0 0 0 0 10 5 ')' • • E z e Q. "0 ~ :> "'"'~ a. 10 4 ItTwo-phas• ~ f low ,egion • Stngle-phase How • Two-phose trow 10 3 1~~~~~--~~~~~~~~~~~~2 3 5 10 20 30 50 100 200 400 Average mog net tempe rolure • K Fig. 5. Pressure drop through 235m of 10.8-mm-ID tube as a function of average temperature (note dramatic change in pressure drop when two-phase ftow is established). 427 428 M. A. Green, W. A. Bums, and J. D. Taylor resulted in a decrease in the pressure drop. The temperature eorresponded to the saturation temperature, whieh was eontrolled by the absolute pressure in the tube. The 2-m test coil behaved similarly to the 1-m test eoils. There were pressure drops during two-phase ftow of less than 104 Pa and no oseillatory behavior. At 4.6 K the CTi 1400 refrigerator supplied two-phase helium to the experiment at the rate of 4g/s. The gas-cooled eleetriealleads performed just as they would have in a bath of helium. Liquid entering the bottom of the gas-cooled Ieads had no effeet on their performanee. Gas ftow through the eleetriealleads was regulated with a needle valve just upstream from a Rotometer-type gas ftow meter. "Garden Hose" Tests The garden hose effeet was tested with aseparate experiment. Detailed data are given by Taylor et al. [13]. Subsequently, the test coil was mounted in series with the 2-m-diameter test coil, in order to test the effeet of pressure drop oseillations on a supercondueting magnet. The test coil was fabricated from 160 turns of 16.6-mm-ID tube. The diameter of eaeh turn was 0.9 m. The tube was divided into two bundles of 80 turns eaeh. Pressure taps and carbon resistor temperature sensors were installed at the ends of the two eireuits and at a eenter tap between the two bundles. Pressure was measured by room-temperature transducers. The test coil was run on the CTi 1400 refrigerator with two-phase helium mass ftows as high as 4 to 5 g/s. The measured pressure drop for this test eoil was about 2 x 104 Pa [2.4 x 104 Pa is predicted using (3)]. In addition, a pressure oseillation with a period of about 30 s was seen. (Short ehops of higher frequeney were seen, but the 30-s period was eonsistently observed.) The peak-to-valley amplitude of this oseillation was around 2 x 104 Pa. Oscillations noted at one pressure tap did not eorrelate weil with oseillations observed at other taps. Temperatures tended to eorrelate with absolute pressure (except at the entry of the experiment). Several Observations ean be made about the experimental data: (1) the pressure drop and the oseillation amplitude deereased as the pressure inereased, the "garden hose" effeet eeased at or near the eritieal point; (2) the oseillation frequency did not vary a great deal with either pressure or mass ftow; (3) other mueh Ionger-period oscillations (about 30 min) were observed when the eontrol dewar was empty and the existence of two-phase ftow was in doubt. HeUum Pump Loop Tests A reciproeating double-aeting bellows-type helium pump was fabrieated and tested at LBL. The pump is driven by a torque motor at room temperature; the reeiproeating motion is transmitted to the pump at 4 K through a shaft whieh operates between room temperature and 4 K. The helium pump is connected to a copper tube heat exehanger whieh has an area of about 2m2 • This heat exehanger removes most of the pump work from the pumped helium stream. The helium pump and its heat exehanger are shown in Fig. 6. More data on the LBL heliumpump are reported by Burns et al. rt 4 ]. The helium pump was first tested in its dewar without any external transfer lines. The volumetrie eftieieney was measured as a funetion of torque motor speed, stroke, and pressure aeross the pump. In general, the highest volumetrie efticiencies were found at the highest speeds and lowest pressure rise across the pump. At a mass ftow of 27 g/s and a pressure rise of about 4 x 104 Pa, volumetrie eftieieneies as high Forced Two-Phase Helium Cooling of Large Superconducting Magnets 429 Fig. 6. LBL heliumpump with tubular heat exchanger. as 80% were observed. At lower pressure rises, volumetric efficiencies of 90% or more were observed. The pump is eapable of delivering about 50 g/s but volumetrie efficiencies were not measured at mass ftows greater than 27 g/s. At low mass ftows the volumetric efficiency dropped below 50%. lt is believed that late closing of one of the popet inlet valves at the lower speeds is responsible for the deeline in volumetrie efficiency. The helium pump was tested in conjunction with the TPC magnet control dewar and transfer lines. The pump was run at mass ftow rates of about 8 and 40 g/s. The eontrol dewar was refrigerated by CTi model1400 refrigerator and the pump was run when there was no refrigeration. At the highest mass ftow (40 g/s), the ealeulated adiabatie efficieney was about 50% . At the lower mass ftow rate (8 g/s), the adiabatie efficieney appeared to drop to about 30%. The measurements are eonsistent with the decline in volumetric efficiency which occurs at low pump mass ftow. CONCLUSIONS The two-phase foree-cooled system is desirable for many kinds of supereondueting magnet systems to be used in aeeelerators, particle physics and in spaee. Two-phase eooling systems are often lower eost than eonventional bath systems. LBL has demonstrated that foreed two-phase eooling works very weil in largediameter thin solenoid magnets. Two-phase ftow systems ean be designed so that they are a reliable way of cooling supereondueting magnets at temperatures 430 M. A. Green, W. A. Bums, and J. D. Taylor approaching that of a bath cryostat. The two key elements in the design of two-phase ftow systems is the elimination of parallel circuits and the minimization of the number of up and down loops which can cause "garden hose" oscillation and pressure drop. ACKNOWLEDGMENTS The authors wish to thank P. Eberhard for his encouragement. They also thank R. R. Ross, H. Van Slyke, and C. Covey for their work on the experiments. This work was performed under the auspices of the U. S. Department of Energy. REFERENCES 1. R. D. McCarty, NBS Tech. Note 631 (1972). 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. M. Morpurgo, CERN, private communication. CERN Courier, "Villigen Superconducting Muon Channel Begins Operation," 15(2):36 (1975). M. MacAshan, Stanford University, private communication. M. A. G,reen, IEEE Trans. Nucl. Sei. NS-18:669 (1971). 0. Baker, Oil Gas J. 53(12):185 (1954). M. A. Green, "Determination of the Safe Operating Point for Hydraulic Operation of the Magnets and Transfer Lines for ESCAR," LBL Engineering Note M4956 (August 1976). R. W. Lockhart and R. D. Marinelli, Chem. Eng. Progr. 45(1):39 (1949). M. A. Green, "MINIMAG Experiment, Large Superconducting Solenoid Magnet, the Cryogenic System." LBL Engineering Note M4834 (June 1975). M. A. Green, in Proc. 7th Intern. Cryogenic Engineering Conference, IPC Science and Technology Press, Guildford, England (1979), p. 86. M. A. Green, "The Development of Large High Current Density Superconducting Solenoids for Use in High Energy Physics Experiments," Doctoral Dissertation, University of California at Berkeley, LBL-5350 (May 1977). M. A. Green, Cryogenics 17(1):17 (1977). J. D. Taylor and M. A. Green, "Garden Hose Test," LBL Physics Note 857 (November 1978). W. A. Bums, "The Construction and Test of a Double Acting Bellows Liquid Helium Pump," tobe published as a LBL report. DISCUSSION Comment by K. D. Timmerhaus, University of Colorado: This paper advocates the application of forced convection boiling for the cooling of superconducting magnets and supports thi& position with operational experience. However, much of the past Iiterature on this type of cooling is less optimistic than the authors. Would the authors care to comment on this difterence in viewpoint? Answer by author: This paper suggests that two-phase cooling is most applicable in devices which are not cryogenically stable. The key to making two-phase cooling work is the control dewar. This reduces the inlet quality of the helium to near zero (allliquid). The two-phase helium is circulated through the Ioad and returned to the control dewar. The only instability we have observed is the so-called "garden hose" instability. The second thing which is important is to make mass ftow per unit area high enough so that the slug-plug ftow region of the Baker diagram is avoided. This requires mass ftows per unit area of greater than 10 kg/m2 • We try to operate our systems at mass ftow rates per unit area of at least 40 kg/m. The skepticism of many relative to this type of cooling is understandable, but I think there is ample experimental evidence to suggest that such fears may be unfounded. The reason is that two-phase helium behaves difterently from two-phase nitrogen or water-stream mixtures. The difterences are that helium is a low-density fluid, the density ratio between the liquid and gas phases is much lower than for other ftuids, and helium has a low critical pressure. H-1 CONTACT HEAT TRANSFER IN SOLID CRYOGENS B. I. Verkin, R. S. Mikhalchenko, V. F. Getmanets, and L. G. Goncharenko Physico-Technical Institute of Low Temperatures Academy of Seiences of The Ukrainian SSR Kharkov, USSR INTRODUCTION Wide-spread use of contact cooling in cold sublimation accumulators employing solid cryogens has necessitated experiments to investigate contact heat exchange with such low-temperature refrigerants [1 ' 2 ]. The first results were not too accurate or reliable and covered only a narrow range of parameters. A thorough analysis of the contact heat transfer has, however, been extremely difficult because of the particular conditions under which it occurs, namely, the relative motion of the heat-exchange surfaces with continuously changing structure of the contact zone and the vapor ftow from the subliming solid-cryogen surface. This work was undertaken, therefore, to obtain more reliable experimental data for solid nitrogen, argon, and methane. lt was also performed to provide a better insight into the contact heat transfer mechanism which occurs during continuous motion of the in vacuo subliming refrigerant. METHOD The apparatus and the technique used in this experimental work have been described elsewhere C]. The advantage of this technique over that described by (involving large systematic errors in the heat transfer coefficient Voroshilov et al. measurement because ofthelarge heat ftux from the outside to the contact zone) is essentially adiabatic conditions achieved in the measuring cell. The latter was a cylindrical vessel, 120 mm in diameter and 200 mm in length, fabricated of 0.3-mmthick stainless steel. The cryogenic liquid under study was solidified in the cell by evacuation of the space above the solid-cryogen sample to a pressure below that of the triple point. Heat was applied through the cell bottom and to the porous piston resting upon the solid-cryogen sample. This piston could be loaded with a maximum of five weights by means of a vacuum maniP-ulator. These weights provided contact pressures of up to 10 kPa (about 0.1 k~/cm 2 ). Based on previous experiments [ ], the experimental technique was further improved so as to obtain more reliable results for the contact heat transfer coefficient. As indicated previously [1], if the heat applied was no larger than 0.1 WI cm2 , the temperature remained uniform throughout the solid-cryogen sample eJ 431 432 8.1. Verlda, R. S. Mikhalchenko, V. F. Getmanets, and L. G. Goncharenko (except for a narrow contact zone, 1 or 2 mm in thickness) and was equal to the saturation temperature associated with the equilibrium vapor pressure. This permitted temperature measurement in the solid cryogen to be made indirectly (by monitorlog only the vapor pressure in the cryogenic refrigerant during the experiment) and thus eliminated the use of thermocouples in the contact zone. In the earlier cell design, many thermocouples and control threads were located in the space near the contact heat exchanger to monitor the motion of the heat transfer zone. Preliminary experiments showed that this prevented good contact between the subliming surface and the contact heat exchanger and thus caused underestimation of the contact heat transfer coefficient [ 1] . The new experiments were also used to further an understanding of the effect that the specific thermal ftux (in the range of q = 10-3 to 10- 1 W /cm 2 ) , the specific contact pressure on the sample (Pc = 0.5 to 10 kPa), the vapor pressure (P = 1 to 500 torr), the position of the heater with respect to the directional vapor ftow and the cryogenic refrigerant temperature have on the contact heat transfer coefficient. EXPERIMENTAL RESULTS The experimental results indicate that the contact heat transfer coefficient is independent of the vapor pressure and also the heater positionrelative to vapor ftow (i.e., ftow from below the solid-cryogen block and through it, or from the top and through the upper heater). The results of the other experiments are shown in Figs. 1 through 4. An analysis of the results suggests that there are two different mechanisms of heat transfer in the low-temperature zone and near the triple point. In order to understand the heat transfer mechanism for solid-cryogen sublimation near the triple point, the authors also studied contact heat transfer during the melting of solid nitrogen and argon (see Fig. 4). 2 0 0 ." . •• ~ ! • B ~ ~ 0 0 0 0 0 ~~ o- I 6 - J 0 - J 2 Fig. 1. Contact temperature difterence 4 T during Sublimation in the low-temperature zone vs. specific thermal ftux q. Conditions: (1),(2),(4), Pc = 0.5-5 kPa ; (3) Pc = 2.5 Pa; (1) methane (T, = 67 K) ; (2), (3) nitrogen (T, = 49--{;1 K) ; (4) argon (T, = 77-82 K). 6 Fig. 2. Eftect of contact pressure on the temperature difterence, 4 T, in the contact zone during solid-argon Sublimation. T, = Conditions : (q = 1.39 W/cm 2 , 78 K): (1) Pc = 5 kPa; (2) Pc = 0.5 kPa; (3) Pc = 10 kPa. Contact Heat Transfer in So6d Cryogens 433 ~1171 '1,.•K Fig. 3. Variation with time, -r, of the contact heat transfer coefficient during solid-argon Sublimation (T, = 78-82 K) for a variety of conditions: (1) q = 1.39 X 10-2 W /cm2 , Pc = 0.5 kPa, P = 30 kPa; (2) q = 1.39 x 10-2 W/cm 2 , Pc = 5 kPa, P =50 kPa; (3) q = 1.08 x 10-2 W/cm2 , Pc = 5 kPa, P = 30 kPa; (4) q = 0.60 x 10-2 W/cm 2 , Pc = 5 kPa, P = 30 kPa; (5) q = 1.78 x 10-2 W /cm2 , Pc = 5 kPa, P = 30 kPa; and (6) q = 1.39 X 10- 2 W /cm2 , Pc = 5 kPa, P = 30 kPa. 174 tll o• c . ooo • 4o• g~•a•J ~ /14 0 "' 0 { o- 2 IJ.P IJ . "- . . . .• - 2 j • "' 0 0 0 0 4 - ~ •-s 0 - 6 6 '{h Consider first the specific features and the mechanism of contact heat transfer in the low-temperature zone. As is evident in Fig. 1, all coolants investigated are characterized by a linear contact temperature drop increase which is dependent on the specific thermal flux. It was also shown by the experiments that the temperature drop in the contact zone is constant for all the cryogenic refrigerants investigated as the contact pressure is increased from 0.5 to 5 kPa. According to previous work C· 2 ], this concept is valid for solid neon, nitrogen, argon and carbon dioxide for contact pressures from 0.5 to 100 kPa. However, additional experiments showed that for solid argon, rapid decrease in the temperature drop (approximately by a factor of 2) occurs as the contact pressure is increased from 5 to 10 kPa (see Fig. 2). This is evidently caused by rapid changes in the contact zone structure as a result of plastic deformation owing to a stress rise in the contact zone to above the solid-argon yield stress. To further elucidate the contact heat transfer under conditions of low contact pressure, the pistonwas replaced with the lightest possible heater, 1.25 g in weight and 50 cm2 in area. The experiments using this modification revealed that a reduction in contact pressure from 500 to 2.5 Pa caused only a relatively small increase (by an order of 1.5 to 1.7) in the contact temperature drop. Even this minor increase seems to stem from heater placement owing to rigidity of the thermocouple and heater Ieads and not from changes in the structure and heat transfer mechanism in the contact zone. These results suggest the conclusion that as the contact pressure is varied from zero to the solid-cryogen yield stress, the contact heat transfer coefficient is constant to a first approximation, irrespective even of the thermal flux (see Fig. 3). The contact heat transfer coefficients in the low-temperature zone for solid nitrogen, argon, and methane are, 0.48 x 10-2 , 0.68 x 10-2 , and 0.75 x 10-2 W /cm2 K, respectively. Fig. 4. Assumed pattern of changes for the contact heat exchange temperature in the vicinity of the triple point and experimental data for the contact temperature difference during sublimation and melting of solid nitrogen. Legend: (1) sublimation at T, = 61.2 K; (2) sublimation at T, = T,; and (3) melting. 434 B.l. Verldu, R. S. Mikhalcheako, V. F. Gebuaets, ud L. G. Gondw'enko DISCUSSION Using these values and other experimental results, one can provide some insights on the mechanism of contact heat transfer when some information on the contact zone structure is known or possibly available from an analysis of experimental data. Estimates show that the minimum contact pressures (2.5 Pa) are much higher than the maximum possible increase in the vapor pressure (-0.01 Pa) acting on the contact heat exchanger. Therefore, there may not be an uninterrupted gas cushion at the solid surface and there must be a number of contact points at the subliming surface. The minimum relative contact area may be estimated from the equality condition that exists between the contact pressure and the sublimingmaterial yield stress, u, i.e., Tim = Pc/u. Thus, for solid nitrogen, with Pc = 100 kPa [2 ] and u = 900 kPa, Tim = 0.1, while for solid argon, Tim= 0.04 (Pc = 7.5 kPa and u = 200 kPa). The remaining area may be occupied by a vapor cushion. As was determined by experiments [1 ], the equivalent thickness of this gas layer may not exceed 1 or 2 mm. The minimum equivalent gas layer thickness may also be estimated because the contact heat exchange can only move in a plane-parallel fashion provided that the sublimation rate and heat ftows are equal over the entire heat exchanger. Then assuming that the heat transfer through the gas cushion is mere thermal conduction, we obtain for the minimum thickness of the equivalent gas layer that ße = Äg/ a. For solid nitrogen, in particular, ße = 0.1 mm. This is the same to a first approximation for the other solid cryogens as weil. Based on the above data for the contact zone structure, one may estimate the contributions of various factors to the contact heat transfer. In the general case, heat transfer across the contact zone may be realized by thermal conduction through the contact points and through the gas cushion by radiation and free or forced convection. With regard to solid cryogens, radiation may be neglected because of the low temperatures used. Free convection in the 0.1- to 1-mm gaps cannot occur because of the low gas density and temperature gradients. To estimate the role of the forced convection owing to gas ftow from the subliming surface and along the heat transfer surface, we need to define the gas ftow behavior. lt is readily apparent that, at pressures above 100 Pa, viscous gas ftow occurs in the gap. For the conditions under consideration, the following relation is readily derived for the turbulent region of forced convection: Re = _g_ 2: 2300 TsJLH (1) where Q is the total heat ftux to the sublimation zone within the perimenter H. The maximum value for the Reynolds number, as obtained by the present experiments (with Q = 10 W, '• = 250 kJ/kg, JL = 6 x 10-7 kg/ms, H = 10 cm) was 100. Thus, for the conditions of contact heat transfer investigated, the gas ftow cannot be turbulent. For effective gas thermal conductivity in the laminar ftow near the subliming surface, Getmanets et al. [3 ] arrived at the relation Ae11 = ÄgNu where the Nusselt number is determined by the following relation [3]: Nu= 1 4TCp) +0.156( 1 +---;;- (2) Contad Heat Transfer in So6d Cryogens 435 Using the experimental value of !l.T for the contact zone (less than 10 K), this gives a Nusselt number value for nitrogen of less than 1.02. Therefore, the forced convection effect on the contact heat transfer may be neglected. Since the longitudinal gas velocity along the heat exchanger is close to the velocity of sound (about 200 m/s), it is desirable to estimate the thermal effects which can arise in the ftow as its velocity is reduced. Proceeding from the law of energy conservation for a ftow, the maximum variation of the specific enthalpy is simply !:ii = V 2 /2 while the largest possible thermal ftux through a unit area is given by !:i.l = !:iim = !:ii!I F F r. (3) For the initial parameters mentioned above, !:i.T/F = 0.5 x 10-2 W/cm 2 , which is 5% of the applied thermal ftux (q = 0.1 W/cm 2 ). The actual effect of reducing the velocity is stilllower because zero velocity and ideal heat transfer for the contact heat exchange cannot be attained. The absence of an effect of the forced gas ftow on contact heat transfer is confirmed by the fact that the latter remained essentially unchanged as the vapor pressure was varied from 0.1 to 50 kPa, proportional to variations in both the transverse inlet velocity and the longitudinal gas velocity along the heat transfer surface. Therefore, the analysis carried out Ieads to the unambiguous conclusion that over the parameter range explored (q = 1 X 10-3 to 3 x 10-2 W/cm 2 , P = 0.1 to 50 kPa, Pc = 0.5 to 10 kPa), contact heat transfer is entirely dependent on the heat transfer through the contact points and the thermal conductivity of the gas cushion. One would expect that the number of contact points on the subliming surface is close to the number of crystalline grains exposed on the surface. This number, as calculated by formulas and using the experimental findings of Verkin et al. [4 ] is 20 cm - 2 for solid nitrogen. This enables one to consider the structure and heat conduction in the contact zone in more detail. Thermal resistance of the contact is determined by two factors: phononthermal resistance and thermal resistance owing to thermal ftux contraction. Estimations using relations developed by Little [5 ] showed that the temperature drop owing to phononthermal resistance at nitrogen temperature is negligible (10-4 K). The thermal resistance is more pronounced owing to contraction of the thermal ftux to n isolated contact points. Under these conditions, according to Shlykov [6 ] a As = 2 71 112 - Jn/Tr 1.41 + 0.371 (4) This relationship, using 71 = 0.1, provides an a value for solid nitrogen of 5.6 x 10-3 W /cm 2 which is very close to the measured value of 4.8 x 10-3 W /cm 2 • With 71 of 0.04, the calculated value of a for solid argon is 2.5 x 10-3 W /cm 2 K (assuming the same number of contact points as for solid nitrogen, n = 20 cm - 2 ) as compared to the experimental value of 6.8 x 10-3 W /cm 2 K. From the relation 71 = nTra 2 we obtain the contact point radius a of 400 ~m for a grain size L of 2 mm. The estimates given above suggest that the thermal resistance owing to thermal ftux contraction is predominant in contact heat transfer. Accordingly, we may approximate the contact heat transfer coefficient in the low-temperature zone from aL; KJ 2AsL:J = (5) 436 B. I. Verldu, R. S. Mlkhalcheuko, V. F. Getmuets, md L. G. Goaeharenko The above mechanism of contact heat transfer prevails at temperatures at least several degrees below the triple point. However, as the operational temperature of the solid coolant ina:,eases, or as the specific thermal ftux increases, the critical condition given by ll.T = (T,- T&)a/q > 1 is no Ionger valid, and the contact heat exchange temperature will rise above the triple point of the coolant. As a result, a liquid layer will appear on the solid surface and the contact heat transfer coefficient will change rapidly; the change is greatly inftuenced by either the boiling or melting heat transfer coefficient. Therefore, this study first attempted to experimentally investigate contact heat transfer near the triple point for both sublimation and melting of the solid cryogen. The experimental work reported here has shown that heat transfer in the contact zone develops as depicted in Fig. 4. If the initial temperature of the solid cryogen is equal to that of the triple point, then as the thermal ftux increases, the contact heat exchange temperature can be represented by curve 9. Thus, just as was the case for contact heat exchange in the low-temperature zone, there isalinear behavior of the temperature drop with thermal ftux growth but with a higher value for the contact heat transfer coefficient (a = 4 x 10-2 W /cm 2 K for nitrogen). The slope of the curve depends on the process underway in the contact zone (melting or evaporation) and the associated heat transfer coefficient. At lower temperatures of the solid cryogen, the process first develops similar to the sublimation curve in lowtemperature zone 4. Thereupon, as the heat exchange reaches the triple point temperature, the temperature drop evolves as one of curves 5 through 8 or parallel to them, as dictated by the initial solid-cryogen temperature. It should be emphasized that the contact heat transfer coefficient for melting is 10 to 15 times as high (4 x 10-2 and 0.1 W /cm 2 K for nitrogen and argon, respectively) as for low-temperature sublimation (4.8 x 10-3 and 6.8 x 10-3 W /cm 2 K, respectively). Thus, the contact heat transfer coefficient for melting is close tothat for Sublimation near the triple point. The rise in contact heat transfer coefficient near the triple point is explained as being due to the liquid layer that develops in the contact zone. The thermal conductivity ofthisliquid layer is at least several times higher than that of the vapor. Therefore, replacement of the vapor by the liquid in the gap near the contact heat exchange Ieads to an enlargement of the effective contact area for each grain of the solid. This effect, in turn, Ieads to a higher heat transfer coefficient. CONCLUSIONS In the course of this work, more accurate values for the coefficient of contact heat transfer to nitrogen, argon, and methane have been obtained for a wide range of conditions. Two distinct laws of heat transfer during solid-cryogen Sublimation in the low-temperature zone and in the vicinity of the triple point have been determined. lt has been shown that in the low-temperature zone there is a critical contact pressure on the solid cryogen (equal to the solid-cryogen yield stress), whose increase results in a sharp change in the heat transfer coefficient magnitude because of plastic ftow at the contact points. The contact heat transfer coefficient for subcritical contact pressure conditions at low temperatures is determined principally by the thermal resistance of isolated contact points and therefore depends on the solid cryogenic refrigerant thermal conductivity and crystalline grain size. Pronounced increase in the contact heat transfer coefficient (by an order of magnitude and more) has been revealed in the case of either sublimation or melting Contact Heat Transfer in Solid Cryogens 437 in the vicinity of the triple point owing to the enlargement of the effective contact area because of liquid layer development in the contact zone. NOTATION a = contact spot radius cp = = = = = L = m= Nu= n = P = Pc = 0 = q = Re = r, = Tc = T, = T, = t. T = V = F H J i gas thermal capacity area perimeter enthalpy specific enthalpy grain size mass flow rate of gas Nusselt number number of contact points per unit area vapor pressure specific contact pressure total heat flux specific thermal flux Reynolds number sublimation heat contact heat exchange temperature solid coolant temperature triple point temperature temperature difference between gap walls, (Tc - T,) gas velocity Greek Symbols a = contact heat transfer coefficient 8, = equivalent thickness of gas layer 11 = contact area 11m = minimum relative contact area A.11 = effective gas thermal conductivity Ag = gas thermal conductivity A, = nonporous cryogen thermal conductivity REFERENCES 1. R. S. Mikhalchenko, V. F. Getmanets, L. G. Goncharenko, and A. V. Polyakov, in Proceedings of6th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1976), p. 293. 2. B. S. Voroshilov, A. B. Grachev, and V. M. Brodyanski, Inzh.-Fiz. Zh. 33:238 (1977). 3. V. F. Getmanets and R. S. Mikhalchenko, in Gidrodinamika i teploobmen v kriogennykh sistemakh, Naukova Dumka, Kiev, U.S.S.R. (1978), p. 24. 4. B. I. Verkin, V. F. Getmanets, and R. S. Mikhalchenko, Cryogenics 19:17 (1979). 5. W. H. Little, Can. J. Phys. 37:334 (1959). 6. Yu. P. Shlykov, E. A. Ganin, and S. N. Tsarevski, in Kontaktnoye termicheskoye soprotivleniye, Moscow, U.S.S.R. (1977), p. 328. H-2 DIGITAL COMPUTER SIMULATION OF VOIDAGE IN A REGENERATOR* J. 8. Harness and P. E. L. Neumannt University of Bradford Bradford, England INTRODUCTION A regenerator is a device which transfers energy between two ftuids by storing that energy in a matrix. Normally the two ftuids are both gaseous and the matrix, in small-scale versions, normally consists of small spheres. Regenerators work over wide temperature ranges and have many applications from the preheaters for blast furnaces down to very small devices found in refrigerators operatingdown to 12 K. This study describes modifications to regenerator theory, developed for the largescale application, so that it may be applied to small-scale applications as weil. REGENERATOR THEORY Regenerator theory has been developed for applications such as Cowper stoves and precooling stages of a tonnage liquefaction plant where the concern is with long-period rapid ftow, and the mass of gas passing in one period is large compared with the amount contained in the void space at any one instant. In order to produce a mathematical model C] of the situation, one of the assumptions made is that at the reversal of the gas stream no gas is left in the regenerator voids. This is not the case with regenerators used in Stirling engines and refrigerators and similar cycles. Here, the ftow alternates in very rapid cycles and only a small proportion of the gas passes through the bed during the course of the cycle; it often happens that this is less than the amount contained in the voids. This is expressed quantitatively by stating that the "holdup," defined as the ratio of the mass of gas contained in the voids tothat passing through in one unidirectional blow period, is greater than unity. If a plug ftow model is adopted where the linear velocity of the gas is regarded as uniform throughout the volume and there is no mixing, it is implied that a certain volume of gas remains permanently resident in the regenerator as illustrated in Fig. 1. In some cycles the holdup is large, while in a Stirling refrigerator values of 0.5 to 1.2% holdup are usually encountered. Another feature of regenerator operation in thermodynamic cycle applications is that the gas ftow rate and inlet temperatures are not uniform throughout the blow period and the physical properties gas and solid are temperature dependent. lt is * Work sponsored by the Ministry of Defence, U.K. t Present address: David Livingstone Consultants Ltd., Brighton, England. 438 Digital Computer Simulation of Voidage in a Regenerator 439 GAS New gas OUT New gas Fig. 1. Effect of voidage in regenerators with short period times. eJ clear from experimental work [2 ] that the conventional regenerator theories are not capable of predicting the performance of regenerators operating in cyclic machines. ASSUMPTIONS FOR COMPUTER MODEL Regenerators normally used are cylinders closely packed with spheres or wire. The gas is channeled along numerous tortuous paths through the bed which are interconnected but have branches at many places and can be regarded as independent of one another. There is no reason for favoring a path at one point in the bed over another, so it is reasonable to adopt a plug ftow model with uniform velocity across the bed. Although mixing on a small scale may occur, it is assumed that longitudinal mixing over a distance large enough to affect the gas temperature profile can be ignored. As in many previous models it is assumed that the thermal conductivity is infinite across the bed while it is zero along the bed. The temperature distribution within the solid may be considered to be uniform. Calculations based on the Hausen [1 ] model indicate that this is the case. Other nonconvective heat transfer processes which may take place in the regenerator (in particular radiation) are neglected. Also initially neglected are the physical property variation with temperature of the gas and solids, although some suggestions are made later as to the development of a more exact representation. In the absence of an effective theoretical analysis of the process of convective heat transfer in packed beds, the heat transfer coefficient must be derived from one of the several empirical correlations available, each of which is intended for a certain range of Reynolds numbers and particle size. The correlations most applicable to the regenerators under consideration and used in this work are those developed by Littman et al. CJ. BASIC EQUATIONS e]. The differential equations which describe the basic processes of heat transfer A heat balance over a small within the regenerator are those given by Hausen element of the regenerator after simplification yields _ m,C, dT, HA, dt = T. _ T. = pAxL,Cp dTg + GCpL, aTg ' g HA, dt HA, ax (1) It is convenient at this point to introduce a reference gas ftow rate Go so that r = G/ Go and a reference heat transfer coefficient Ho so that (J = H/ Ho. Equation J. B. a . - ud P. E. L Neuwut 440 (1) can be rewritten using the following dimensionless quantities: = HoA,x/ GoCpL,., 71 = HoA,J/ m,C,., q, = pAxL,Cp/ m,C,., Ao = HoA,/ GoC", 1T = HoA,T/ mrC, lt should be noted that the dimensionless void fraction q, represents the ratio of ~ the heat capacity of the gas in the void to that of the solid matrix. The parameters Ao and 1To are the Hausen reduced length and period applied to a regenerator Operating under constant ftow rate G 0 • Under these conditions Hu Ao = q,- (2) 1To Thus, equation (1) becomes _ aT. = 8 r(T. _ T. ) = cf> aTg + raTg a71 • a71 g where both 8 and rare functions of Go. For constant ftow rate conditions r thus (2) becomes _ aT. = T. _ T. = -~.. iJTg + iJTg a71 • g o/ (3) a~ a71 a~ = 8 = 1; (4) which is similar to that used by other workers [4 ' 5 ]. The boundary conditions are determined at the reversal and inlet conditions. For each ftow period the initial values of both the solid and gas temperatures at all points in the bed are required, but in counterftow operation at the beginning of a period, these are placed equal to the values at the end of the previous period, i.e., T; (~. 71 = 0) = T~ (A- ~. 71 = 1T) T~ (~, 71- O) = T~(A- ~. 71 = 1T) (5a) (Sb) The gas inlet temperature is also a function of the condition (6) Equation (3) differs from the simpler theories [1 ' 6 ] only in the presence of the extra term involving q, but a coordinate transformation can be used to eliminate this term, namely, z = ~. y = 71 - q,~, or inversely ~ = z and 71 = y + cf>z. With this transformation equation (3) becomes = T. _ Tg = iJTgl _ aT., ay az z (7) y The usual finite difference technique [6 ] may be used to solve these equations. Certain modifications to this procedure are necessary at the boundaries of the mesh created by the technique because of the change in the coordinate system; the boundary conditions have to be changed to eJ T; (z, y T~ (z, = -t/>z) = y = -t/>z) = T; (A- z, y T~(A- z, = 1T- t/>z) (8a) y = 1T- t/>z) (Sb) 441 Digital Computer Simulation of Voidage in a Regenerator t t z z A .J-. -VA '~ ... "' .J-~ . TI-VA ~ 0 ~ 0 TI Y- -VA TI TI-VA Y- (b) (o) Fig. 2. Effect of transformation on the finite difference grid. (a) Where the reversal coincides with the grid points; (b) where the reversal does not coincide with the grid points. The effect that these transformations have on the solution is illustrated in Fig. 2. lt should be noted that the rectilinear coordinates y and z are those in which (7) is defined; hence, in the finite difference technique [6 ] which is to be employed, a rectilinear network must be superimposed. Figure 2 shows how this is achieved. Commencing with the initial conditions specified on the left-hand sloping boundary, it is possible to advance the solution, working in rows parallel to this edge, through the entire network. There are problems, neglected by previous authors [4 ' 5 ], in this method because of the Iack of freedom in closing the grid ratio, namely, that the increments are related by the expression Äy = <P Az (9) If batch Ay and Az are to be reasonably sized integral fractions of A and 7T, respectively, it is necessary that the quantity Hu = <PA/ 7T be a rational number with small numerator and denominator. If one does not adhere to this imposition, reconciliation is best achieved by interpolating the conditions at the end of the period from the calculated grid on either side. Here the variations in temperature are likely to be at their smallest in the time abscissa y. Another expedient available when Hu < 1, and particularly useful for small </J, is to change the time increment Ay over part of the range. This permits economies in the number of time steps and hence computer time. lt is also possible when solving (7) to make use of a more rectilinear grid; hence the two step sizes may be chosen independently, each satisfying its own error condition, while the number of steps can be maintained at a reasonable Ievel. This technique is basically unchanged from the previous method, but generates a parallelogram network of elements as in Fig. 2b. Any formula for interpolation which makes use of the available information is acceptable but must be consistent with the order of the truncation errors inherent in the trapezoidal formula for the differential equations; a simple linear expression is normally adequate. The second of the above systemswill be referred to as the interpolation method and contains a fault which has been disregarded by previous workers. TUE GAS-GAS INTERFACE At the onset of ftow reversal, gas commences to enter the regenerator through the previous exit port, pushing the gas contained in the voids through the bed at reversal. At the boundary where these two bodies of gas meet there is a discontinuity in temperature. In reality, diffusion and conduction processes would soon eliminate this jump in temperature, but this mathematical model contains no terms to account for these effects. These discontinuities create problems with the finite step deviation 442 J. B. Hamess and P. E. L. Newnann approximations, which, dependent as they are on the absence of high-order deviations, break down under such circumstances. This would occur with any derivative along a path which crosses the discontinuity. In fact, the gas temperature derivative aTg/ax is taken along the constant y line, which represents a characteristic of the ftow; in other words, it is taken along a path tracing the history of a gas particle as it passes through the bed. Thus, it is proportional to the substantial or proper time derivative DTn/ Dt. These characteristic lines include the path traveled by the discontinuity; so the gas temperature derivatives run parallel to and do not cross it. The solid temperature derivative does not cross the boundary, but here the temperature discontinuity is only of the second order and much smaller than that of the gas because of the higher heat capacity. The situation may be properly dealt with if two gas temperatures are retained at all points on the interface, i.e., the line y = 0 in Fig. 2. The old gas temperature is used in all expressions relating to the old gas region and the new gas temperature. The resulting boundary profile may be calculated without any reference to the old gas region. This is possible with the reetangular grid technique of Fig. 2a, but not the interpolation method of Fig. 2b because of the necessity to perform an unacceptable interpolation across the discontinuity. lt is found that the two boundary gas temperatures rapidly approach one another, the differing values for the solid/gas temperature drops leading to different rates of heat transfer which tend to achieve this effect. In many cases, by the time the gas exit is reached, there is no perceptible jump in temperature. A comparison of the results of computer programs using the two methods reveals that, although in the body of the regenerator there are some discrepancies, the exit gas curve and the effectiveness computed from this curve are the same. For the interpolation method the ambiguous initial gas entrance temperature (at the point x = 0, y = 0) is set to the mean of the two alternative values, namely, the final exit temperature from the previous period and the initial inlet temperature for the current period. This simulates some form of diffusion, although no attempt is made to account for this in detail. MODIFICATIONS FOR VARIABLE GAS FLOW Variable gas ftow is introduced differently in the two methods. In the interpolation model each strip extending one step in the y direction and running parallel with the sloping initial ordinate, represents a one-time interval in the action of the complete regenerator. Successive strips can be regarded as a sequence of short, individual cocurrent blow periods. To each of these may be assigned a different ftow rate, set as the reference value Goforthat strip. The associated quantities fJ and rare then introduced so that (7) can be solved with values appropriate tothat strip. Thus, a continuously varying ftow rate is represented by discrete values; each value is the average over a single time step. This is an approximation consistent with the numerical procedure. In the reetangular grid model the situation is complicated by the compulsory connection between the step sizes with the constants involved varying with ftow. A better technique is to use a different transform from above, devised directly from (3), namely, aT. aT. (lOa) -• = !1(y, z)-• a71 ay ..~.. aTg _ 1 ( )aTg r-aTg+ ' 1 ' - - y,za~ a71 2 az (lOb) Digital Computer Simulation of Voidage in a Regenerator 443 where ft and [2 are functions tobe defined later. This Ieads to the requirements The inverse relationships can be derived and after lengthy calculations give the results ( az) a~ 1J = !2 (lla) f(7J) ( az) = 0 a1J e ( ay) <P/1 a~ 1J = (llb) (11c) f(7J) ( ay) = !1 a1J e (lld) The first two of the relationships in (11) imply that h = f(17 )/(~), while substitution of (10) into (3) indicates that it is desirable to have / 1 = [2. The simplest choice is for /1 = h = r, so that (11) becomes (az) =1 a~ 1J (12a) (12b) (ay) a~ 1J = -<P (12c) (12d) Integration then gives (13a) z=~ y = L1J f(17') d17'- cP~ (13b) Thus, equation (3) in the new Coordinates becomes aTs fJ aTg --=-(T -T)=- a1J r s g az (14) This is the same as (7) apart from the presence of the factor 8/f which enters the finite difference scheme although it will not alter the distribution of the points on the grid. As this factor is a function of time, it Ieads to varying values of the multipliers in the numerical scheme as one calculates the grids. The definition of y, incorporating the integration of the ftow term, gives precisely the transformed variable required for the evaluation of the substantial derivatives of the solid and gas temperatures. lt J. B. Raraeis ud P. E. L. Neuwm 444 may be regarded as an advancement coordinate, mapping intervals in the progression of the gas front regardless of any irregularity in the elapsed time coordinate. VARIATIONS IN PHYSICAL PROPERTIES WITH TEMPERATURE AND FLOW Many of the physical properties of the gas and solid that appear in these equations are temperature dependent; this renders the solution nonlinear and not directly amenable to the methods discussed above. Willmott proposed an iterative/perturbation scheme which may be incorporated into the interpolation method set forth above with the grid network in the latter maps properly applied onto Willmott's grid as a whole. The reetangular grid transformation model cannot be modified in such a manner. An attempt to introduce the necessary and unpredictable variations in the dimensionless step lengths at each grid point Ieads to a complete description of the transformation system. Under theseadditional requirements, only one of the models is now available. A vastly increased amount of computation is involved and the need to adopt an iterative technique destroys the essentially linear nature of the solution. n CONCLUSIONS The methods discussed above were programmed in FORTRAN for the ICL 1904S computer at the University of Bradford. From the results it is clear that if one uses the simpler design method, such as those of Schalkwijk [8 ], one obtains an answer for the effectiveness of the regeneratorthat is smaller than that attained from the methods discussed above. This is illustrated in Fig. 3. 1.0 Iffi 0.9 Fig. 3. Variation of effcctiveness with reduced length. Digital Computer Simulation of Voidage iu a Regenerator 445 NOMENCLATURE A, Ax CP C, G G0 H H0 Hu L, m, Tg T. t x y z = total surface area of packing for whole bed = area of hydraulic x section = heat capacity of gas = heat capacity of regenerator packing = gas ftow rate = reference gas ftow rate = heat transfer coefficient = reference heat transfer coefficient = holdup = length of regenerator = mass of packing of regenerator = temperature of gas = temperature of regenerator packing =time = distance along the regenerator = dimensionless time, new transform = dimensionless length, new transform Greek Symbols ~ = 11 = p = T = q, = A= 1T = dimensionless length dimensionless time density of gas period time of regenerator ratio of heat capacities of gas in voids and matrix reduced length reduced period Superscripts ' = bot period " = cold period REFERENCES 1. H. Hausen, Wärmeübenragung in Gegenstrom, Gleichstrom, und Kreuzstrom, Springer-Verlag, Berlin, West Germany (1950). 2. A. Bretherton, W. H. Granville, and J. B. Harness, in Advances in Cryogenic Engineering, Vol. 16, Plenum Press, New York (1970), p. 333. 3. H. Littman, R. G. Barvile, and A. H. Pulsifer, IEC Fund. 7:554 (1968). 4. P. L. Heggs and K. J. Carpenter, Trans. Inst. Chem. Eng. 54:232 (1976). 5. A. J. Willmott and C. Hinchcliffe, lnt. J. Heat Mass Transfer 19:821 (1976). 6. A. J. Willmott, lnt. J. Heat Mass Transfer 7:1291 (1964). 7. A. J. Willmott, Int. J. Heat Mass Transfer 11:1105 (1968). 8. W. F. Schalkwijk, Trans. ASME Series A, J. Eng. Power 81:142 (1959). B-3 SIMULATION OF COOLDOWN UNDERNEATH LARGE CRYOGENIC STORAGE TANKS M. H. Seeland* and K. D. Timmerhaus University of Colorado Boulder, Colorado INTRODUCTION With worldwide energy demand continuing to expand and corresponding petroleum supplies questionable, natural gas demand is expanding rapidly. The trend toward higher natural gas consumption is being observed throughout the world. As a result of the increasing demand and higher prices, gas reserves have increased by development of existing gas fields and with new discoveries. According to Hiwada C], confirmed reserves are currently about 2200 Tcf, but undiscovered reserves are estimated at 7500 Tcf. Many years of gas production could be maintained by these confirmed and undiscovered reserves, even if the current 50Td/year production were to double to 100 Tcf/year, the energy equivalent of current worldwide oil consumption. Major energy-consuming countries, such as the Western European nations and Japan, do not have abundant reserves and therefore depend on importing natural gas. Transportation, therefore, has become one of the major problems in natural gas expansion. Naturalgas is presently transported either by pipeline or by tanker in the form of liquefied natural gas (LNG). Unfortunately, either transportation method requires huge investments and long Iead times. Actual output has, therefore, lagged appreciably in most gas-producing countries and predictions of future demand and output are highly questionable particularly in view of the magnitude of the increases (although most forecasts show major increases). The mode of international trade in natural gas will be changing since projected LNG imports are expected to account for 60% of the total natural gas imports. The present situation has about 80% of the natural gas imports handled by pipeline. Even though LNG trade is now fifteen years old, progress has fallen short of expectations owing to the need for guarantees for the huge investments. Currently, however, the increased demand and the improved prices have established a new era of cooperation and progress in the LNG field. Various international projects are being planned now for implementation in the early 1980's. Storage is important in an LNG system for both base Ioad and peak shaving plants. Above-ground double metal-wall tanks are the most common type of LNG * Present address: IBM, Boulder, Colorado. 446 Simulation of Cooldown Undemeath Large Cryogenic Storage Tanks 447 storage tanks. The inner wall of such a storage tank is constructed of a steel alloy (e.g., 9% nicket steel) exhibiting acceptable properties at cryogenic temperatures, while the outer wall is generally constructed of mild carbon steel. Since such mild carbon steels become brittle below -50 to -100°C and frost heaving can occur if the soil below the tank freezes, a potential hazard to the tank is created if cold temperatures progress too far from the inner tank. Therefore, the overall insulation scheme and heater design below such tanks are extremely important. Failure of the outer tankwallnot only reduces the generat tank support, but also results in methane leakage since the insulated space between the walls, ftoor, and roof of the tank is normally saturated with methane gas. Design optimization of an above-ground double metal-wall storage tank includes many factors. The more important factors are (1) insulation material and thickness, (2) location and spacing of the heaters below the tank,. (3) heat input to the heaters, (4) location of the heater temperature control points, (5) minimization of heat leak to the tank, and (6) power failure backup energy systems. These factors must be considered simultaneously to minimize material cost and heat leak to the tank while preventing the possibility of failures owing to low-temperature embrittlement of the outer tank wall and frost heaving underneath the tank. To aid in the analysis of these many design factors, a simplified computer simulation of the transient two-dimensional heat transfer occurring underneath such a tank was developed in this study. The program simulates temperature proflies below and around the tank as a function of time. Generalization of the program was undertaken in order to facilitate the theoretical analysis of numerous design options in a relatively short time. MATHEMATICAL MODEL The development of a mathematical model useful for simulating the transient two-dimensional heat transfer occurring underneath a cryogenic tank durlog cooldown must recognize numerous factors. For example, several different materials will generally be involved; often with quite complicated geometries. Heat input sources are involved and these are controlled and energized in a variety of ways. The mathematical simulation model can have both insulated or constant temperature boundaries depending upon the situation at band. For example, an insulated boundary could be specified to provide an approximation for a boundary that is considerably removed from areas of large heat ftux under the cryogenic tank, while a constant temperature boundary could be the inner ftoor of the cryogenic tank once liquid covered the ftoor. The equations for such a transient two-dimensional conductive heat transfer simulation can be developed from a generat energy balance equation presented by Bird et al. e], namely, a ae (pCvT) = -(V·q)- P(V·v)- (1':Vv) (1) where q is the heat ftux vector, v is the velocity vector of the system, and 1' is the shear stress tensor. For the specific case being considered, the velocity vector is zero and Cv is equivalent to Cp. With the addition of a heat source term, the relation simplifies to aT c -ae = -(V·q) + o P P (2) M. H. Seelud aad K. D. 'I1IIuaerlwls 448 Using Fourier's law of heat conduction (q = -kVT) and assuming that the thermal conductivity of the material is constant, (2) can be expanded in terms of reetangular coordinates to a2 T a2 T ar pC-=k-+k-+Q 2 2 p ax ae (3) ay Equation (3) is the transient two-dimensional heat transfer equation that was used to develop the numerical approximations for this simulation. Although the steady-state form of (3) can be solved analytically, the transient form realistically requires the use of a computer. In order toreformwate (3) into a form suitable for computer analysis, either finite-element or finite-difference numerical approximations could be used. Both methods establish a grid of points in the area of solution. The finite-element method integrates the equation over finite differential areas. These areas are usually triangles in order to allow for improved curve-fitting capabilities. The finite-element method is usually applied to microscopic or simple system analysis owing to the large computer time requirements associated with the integration routines. The finite-difference method, on the other band, estimates the differential equations with numerical approximations and, therefore, requires no integration. The numerical approximations are based on changing the differential variables to discrete variables. For example, the exact relation to calculate temperature 7f+ 1 located at x + Llx relative to temperature 1j located at x in a transient heat conduction simulation is given by a Taylor series as )(;Pr) (a )(a-axzT) i+ (a__K_ + ··· + ax i 3! 3 2 2 ( ar) + ~ Ti+I=1f+Llx2! ax i -3 (4) Upon rearranging (4), the firstderivative takes the form ( aT) ax i = 7f+ 1 - Llx Tj _ (ax)(a 2 ~) 2! ax i _ (ax 2)(a3 ~) ax 3! i _ ••• _ (5) The second derivative takes the form (aaxT) i = 2 2 7f+l- 2Ti + 1f-1 (Llx) 2 _ (ax 2)(a4 ~) 12 ax _ .•• _ i (6) In approximating equation (3J with the finite-difference method, similar equations can be developed for (a 2 Tjay ) and (aTja8). With the aid of these relationships, {3) becomes pCp(T~t~; T~i) = k [Ti,j+l -(!;)~ + Ti,i-1] + k [Ti+l,i -(!~~~ + Ti-l,i] + Q (7) where the temperatures on the right-hand side of the resulting equation can be either at the previous time interval n {explicit) or the new time interval n + 1 (implicit). Upon rearranging the explicit form of the temperatures in (7), the new time interval Simulation of Cooldown Underneath Large Cryogenic Storage Tanks 449 can be calculated from rzt = (aäll) [ (aäli)TZj + .:l~2 (TZj-1 - 2TZi + TZi+d + .:l> (T7-1.i- 2TZi + T7+1,j) + ~] (8) where a = kj(pCp). Equation (8) is simple to formulate and easy to use, but, unfortunately, explicit solutions have restrictions. In order to assume convergence to the differential equation solution as Llx, äy, and Llll are decreased, the ratios Llll/ Llx 2 and Llll/ ä/ are limited in magnitude. The convergence criteria for (8) is [3 ] 1 Llll (9) ~ 2 [(.:lx) 2 + (äy) 2] Therefore, even though Llll could be increased as steady state is approached, the convergence criteria prescribed by (9) may limit the size of Llll. The implicit form of (7) results in an equation with five unknowns. lts use would result in the need to solve a large matrix for the temperatures at all points for the new time interval. In order to establish smaller matrices for solution, an implicit method, the alternating-direction implicit (ADI) method 4 ], was finally chosen for the simulation calculation. Since the ADI method is unconditionally stable, it requires no convergence criteria. The ADI method uses two different finite-difference solutions for each time interval. An intermediate time interval of (ll + .:lll/2) is used to determine intermediate temperatures of T*. These intermediate temperatures are then used to solve for temperatures at time (ll + Llll). The first finite-difference equation is developed from (7) (explicitly in the x direction and implicitly in the y direction) and involves only three unknowns. This permits establishment of a tridiagonal matrix for each column of temperatures (i.e., for each j, a matrix is set up to solve for T* at all i's). The second finite-difference equation of the ADI method is developed from (7) explicitly in the direction (T*'s are known) and implicitly in the x direction. This permits the establishment of another readily solvable tridiagonal matrix for each row of temperatures (i.e., for each i, a matrix is set up to solve for Tatall j's). The two basic equations for the ADI method upon simplification are as follows: e· y -T1-1.i + 2(Ay + l)T[i- .:ly2) [T;,j-1 = ( .:lx2 + 2(Ax -T;,j-1 T1+1.i (lOa) (0) k (lOb) + 2(Ax + l)T;,i- T;.i+1 äx 2) [T;-1.i * + 2(Ay= ( .:ly 2 (Q) k + T;.i+1] + Lly 2 - l)T;,i * l)T;.i * ] + Llx 2 + T;+1.i where Ax = äx 2 /allll and Ay = Lly 2 /aä8. These relations, such as (lOa), can be rewritten for computer use in a generat form as (11) M. H. Seeland and K. D. TimiDerbaus 450 where constants A, B, and C are functions of p, C"' k, and äy, and constant D is a ~unction of p, C"' k, äx, äy, Q, and T~i-1• T~i> and T~i+l· As noted previously, a simulation generally involves many types of materials with widely differing properties and geometric shapes. Thus, there can be points in the established grid that do not conform to (1 Oa) and (1 Ob). Included in the grid could be interfaces between materials, corners of material interfaces, additional heat sources, insulated boundaries, and constant temperature boundaries. These special cases require further modifications of (lOa) and (lOb) before they can be included in the final computer program. Their development is rather lengthy but straighttorward and will not be detailed herein. The inclusion of natural convective boundaries in the Simulation model also requires some reformulation of the basic equations before they can be used in the simulation model. A heat balance along a boundary subjected to natural convection can be expressed by k(~~) = h(Ts- Tamb) (12) where T. and Tamb are the surface and ambient temperatures, respectively. The convective boundary temperatures will change very gradually owing to the assumed constant ambient conditions. Because of these small changes, (12) can be solved explicitly for all points along the convective boundaries and these temperatures are considered as constant for the AD I method solution. The resulting amplified relation for convective heat transfer from horizontal surfaces developed for this simulation model is as follows: Tn+l I,J = 2aMJh) rn. + }:_ Tn (t _I__ Ay äyk Ay <,J . + (2aä8h) T amb äyk <+l,J (l 3 ) A similar relation has been developed for the convective heat transfer from vertical surfaces. As noted earlier, convergence criteria must be satisfied with explicit formulation. For the convective heat transfer relations, the convergence criteria as developed by Adams and Rogers [5 ] are 1 aä8 --<---äy2- 2 + häy/k (14) This sets the maximum ä8 as (15) In order to predict the temperature of the points on the convective boundary for the ä8 interval of the implicit formulation, the explicit equation must be repeated for many ä8max increments. The nurober of repetitions necessary is given by the ratio of ä8/ä8max• As discussed previously, the ADI method used in the program establishes a matrix for each column of temperatures for the first set of calculations and each row for the second set of calculations. Each column or row begins and ends with either a constant temperature or insulated boundary. The resulting equations set up a tridiagonal matrix of the form Simulation of Cooldown Undemeath Large Cryogenic Storage Tanks B1 C1 A2 B2 0 A3 0 D1 D2 D3 0 c2 B3 c3 0 A, B, c, 0 0 0 AR BR D, 451 (16) DR The matrix coefficients (A, B, C), as developed from relationssuch as (10) and (11), are calculated for allsolid materials, horizontal interfaces, vertical interfaces, and corners. These coefficients are recalculated only when the time increment changes. Since the D coefficients are functions of the previous temperatures, they are recalculated for every time increment. Solution of the resulting matrix uses Gaussian elimination (i.e., the Thomas algorithm). A heat source can be simulated in the program either at a material interface or imbedded in one of the materials. Such a heat source can be controlled as desired to provide energy input to the transient heat transfer system under study. This is accomplished by locating a control point for each source and monitaring the temperature at this point after each time increment. If the temperature at the control point falls below a specified temperature, the program energizes the heater system and the designated energy is added to the appropriate ADI metbad equations for the next time increment calculations. This energy input will continue to be added until the control point temperature attains an upper specified temperature. The control point location and desired temperature range for this control point can be input as desired. The final option for heat sources is that a time can be specified, whereupon the energy input from the heat sources can be set equal to zero. This is essentially a "shutdown" time for the heat source and represents apower failure forareal storage tank. Any number of time loops can be set in the program. Foreach time loop, !:18 and the ending time for the loop are set. Since the program is unconditionally stable, solution feasibility is independent of the choice of !:18. But, since error is a function of !:18, trials have to be made to insure that !:18 is sufficiently small to provide consistent results. The error associated with the time increment approaches zero as steady state is approached; therefore, time increments can usually be increased as time progresses. However, when there are changes being made to the temperature distribution, the time increment must be decreased until steady state again is being approached. In order to calculate overall heat lass or gain at a boundary, the temperature differential at the boundary is required. The option to have the program average temperatures is available. Using the surface temperature averaging option, the average temperature differential and, therefore, the average heat flux can be determined at any boundary. Theseaverage surface temperatures arenot only an average of all temperatures along the boundary, but also are an average over time from time zero until the specified end of the program. DISCUSSION OF RESULTS The computerprogram that has been developed in this study was generalized to provide substantial flexibility in evaluating time-temperature effects of design M. H. Seeland and K. D. Timmerluaus 452 Power Off 30 20 u 0 10 ,j ... ...." .e . ..."' "'- 0 -10 '""" -20 -30 3000 0 4000 Hours Fig. 1. Temperature vs. time at three specific locations underneath a well-insulated cryogenic tank supplied with heating coils. Legend: 1, location near center of tank but at some distance from heating coil; 2, location adjacent to heating coil; 3, Iocation near edge of tank but at some distance from heating coil. changes in specific cryogenic storage tanks. In addition, the generality of tht equations allows the program to be useful in many other heat transfer applications. Essentially any case that concerns heat conduction can be readily simulated with the program. Minor modifications can be made to handle more detailed convection calculations or boundaries other than the constant temperature and the insulated boundary situation considered here. Owing to the input ftexibilities, essentially any geometric shape can be simulated. The program currently requires 90° corners, but could be modified to handle irregular boundaries. Heat sources and constant temperature interfaces are already incorporated in the model and these options can be used to simulate heating or refrigeration coils if needed. The computer program was originally used to predict the time-temperature relation at various specific locations underneath a well-insulated cryogenic tank supplied with a heat coil controlled from one of these locations. Figure 1 presents a time-temperature profile for three of these locations. One of the locations was near the center of the tank at some distance from the heating coil. The second location was adjacent to the heating coil and was used as the control point for the heater. The heater was energized when the temperature at this location decreased to 5°C and was programmed for shutoff when the temperature at the location reached l5°C. The third location was near the edge of the tank but also at some distance from the heating coil. A study of the three selected time-temperature curves indicates that in the first 900 hr of cooldown the heater was not energized because the control point temperature was above soc during that time period. Once this specific temperature was attained, the heater was energized for approximately 300 hr with a resultant increase in temperature at this location to l5°C. The beater was then deenergized and the cooling process repeated until the temperature bad again dropperl to 5°C, where- Simulation of Cooldown Undemeath Large Cryogenic Storage Tanks 453 upon the entire process was repeated. The other two locations which were at some distance from the heater did not show the large changes in temperature that were characteristic of the control point. The effect of the periodic heater Operation at these locations was observable but was also fairly weil damped out. At the end of 3000 hr of operation, the heater was completely shut oft to simulate apower failure. Since there was no periodic energy makeup by the heater, the temperatures at all three locations showed continued decreases in temperature for the next 1000 hr of monitoring. At the end ofthistime it was evident from the time-temperature record of all three locations that steady state still bad not been attained. This is not surprising since 1000 hr (-40 days) is still a relatively short time in the cooldown history of many cryogenic storage tanks. More recently the computer program has been utilized to follow the timetemperature history for a large LNG storage tank. In this task it has been quite successful in matehing temperatures observed at numerous locations over extended time periods and under varying heater inputs. Differences between observed and predicted temperatures have been on the order of 1 to 2°C. The principal disadvantage to the program has been its rather rigid grid structure, which provides some difficulty in establishing temperature profiles along each material interface encountered underneath a typical cryogenic storage tank. Present studies are attempting to minimize this deficiency of the simulation model. CONCLUSIONS In general, the design of insulation and heater systems to protect cryogenic storage tanks is complicated, but extremely important. A common pitfall is to design these systems, analyzing each part separately, and then failing tobring these parts together to establish the complete design. With so many interrelated parts, it is important that these be carefully integrated into the overall system. Any simulation program similar to the one developed in this study can be a great aid in making such an analysis. NOTATION CP = c. = h = k = P = q= Q = T = T~ 1 = T!':i = v= x= y = heat capacity at constant pressure heat capacity at constant volume heat transfer coefficient thermal conductivity pressure heat input heat input in heat transfer equations temperature temperature at point (i, j) at time n intermediate time temperature at point (i, j) velocity vector distance in x direction distance in y direction Greek Symbols p = density a = thermal diffusivity T = shear stress tensor 8 =time 454 M. H. Seeland and K. D. Tunmerhaus REFERENCES 1. R. Hiwada, CEER 10 (11):9 (1978). 2. R. B. Bird, W. E. Stewart, and E. N. Lightfoot, Transport Phenomena, J. Wiley and Sons, Inc., New York (1966) p. 314. 3. B. Carnahan, H. A. Luther,-and J. 0. Wilkes, Applied Numerical Methods, John Wiley and Sons, Inc., New York (1969), p. 429. 4. J. Douglas, Jr., J. Soc. Ind. Appl. Math. 3:42 (1955). 5. J. A. Adamsand D. F. Rogers, Computer-Aided Heat Transfer Analysis, McGraw-Hill Book Co., New York (1973), p. 181. H-4 TRANSIENT POOL BOILING OF LIQUID HELIUM USING A TEMPERATURE-CONTROLLED HEATER SURFACE P. J. Giarratano and N. V. Frederick NBS Thermophysical Properlies Division Boulder, Colorado INTRODUCTION Many superconducting devices are subject to transient heating. These transients occur during normal Operation in pulsed magnet energy storage systems, or they may occur in superconducting magnets because of ftux jumping or mechanical instabilities in the magnet windings. The thermal stability of the superconductors under these conditions depends upon the effectiveness of the coolant (helium) to absorb the heat and cool the superconductor below its critical temperature so that anormalzonewill not be propagated. Because the time scale for the heating, resulting from the above types of disturbances, is of the order of milliseconds, the recovery process depends on the transient heat transfer rate. Therefore, efficient cryostability analysis relies on a knowledge of the transient heat transfer characteristics of liquid helium since steady-state values may lead to unnecessarily conservative designs. For pool-boiling systems, cryostability analysis requires information on the entire q vs. !l.T curve including the nucleate-boiling, transition-boiling, and filmboiling regimes. A Iiterature search indicates that the transition-boiling portion of the curve has not been established for helium even under steady-state conditions. This is owing primarily to the hysteretic effect in going from the nucleate- to film-boiling regime and the fact that the controlled parameter in the experiments has traditionally been the heat ftux rather than the temperature difference. Transient helium heat transfer data have previously been reported by Jackson [~, Tsukamoto and Kobayashi ~2 ], Bailey [3 ], Iwasa and Apgar [4 ], Steward [ ], Schmidt [6 ], and Brodie et al. [ ]. In all these sturlies heat ftux was the controlled parameter. In this experiment, a specially designed electronic temperature conteoller has been employed to linearly vary the temperature vs. time of a carbon film heater surface submerged in a liquid helium bath. The entire boiling curve q vs. ll.T, including transition boiling, is obtained for various values of dT/ dt ranging from 0 (steady state) to approximately 7500 K/s. 455 P. G. Giarnt.o ud N. V. Fredericl 456 Carbon Film Surface f 10.5 cm • 0.334 cm l Carbon Film Surlace 2 lnsula1ion 1.3 cm 0.1cm 0.64cm T Ouar1z '- ~ ~ Carbon film (0.5pm lhickl : : ILialion Fig. 1. Description of test sample. EXPERIMENTAL SYSTEM Test Sampie The heater surface, which also served as the thermometer for determining the surface temperature, is a 0.5-JLm (5000-A) thick carbon film which was vapor deposited on a 1-mm-thick quartz substrate (see Fig. 1). The estimated thermal diffusion time across the film is of the order of nanoseconds. The surface area of carbon film surface 1 is 0.167 cm 2 (0.5 x 0.334 cm). This particular sample was used in a previous experiment which required adjacent heater surfaces and this is the reason for the two carbon films. However, in this experiment only the smaller surface was utilized and calculations have determined that lateral heat transfer was negligible. In the earlier part of the transient heating, where molecular conduction predominates, the ratio of the heat flux going into the helium to that going into the quartz substrate is approximately (1) For helium and a quartz substrate at 4 K this ratio is approximately 16. In the steady state the heat flux ratio is ~. ~i s I 1) q,/q. = h, ( -k + -k- + -hif (2) where h1 is the film coefficient between the carbon film and fluid, h;1 is the film coefficient between the insulation and fluid, and ~.. ~; are the thickness of the substrate and insulation, respectively. Using a range, 0.05 s h1 s 0.8, and ignoring 1I h;1 since it is small compared to ~;/ k;, one obtains 50 < q1/ q. < 800. Therefore, in both the transient and steady state, the heat transfer is largely unidirectional from the heated surface to the fluid as desired. The sample is suspended in a vertical orientation in a 33-liter liquid helium storage dewar which is vented to the Boulder, Colorado atmospheric pressure of 0.82 atm (8.3o9 x 104 Pa). Therefore, the data were obtained under saturated conditions with bulkfluid temperature of 4.02 K. Transient Pool Boiling of LHe Using a Temperature-Controlled Heater Surface 457 VR!ll 1 From Function Generator Rlest RSTD Dual ChaMel Digital Record ing Oscilloscope Fig. 2. Schematic of electronic temperature controller and data acquisition equipment. Tempersture Controller The essential feature of this experiment was sweeping the temperature of the carbon film linearly with time. This was accomplished by an electronic feedback control system, the basic components of which are shown in Fig. 2. The output of the operational amplifier controls the power supply current through the test sample so that the output of the analog voltage divider (which is proportional to the test film resistance) matches that of the VR(T), drive signal. VR(T), was predetermined to give the desired temperature vs. time characteristic. For the steady-state data, the drive signal VR(t) was a steady-state voltagerather than a time-sweeping voltage. Although not shown in the schematic of Fig. 2, it was also possible to control the power (heat flux) rather than the resistance (temperature). The steady-state data obtained were the same under both modes of operation. Data Acquisition The voltage drops across the carbon film and a standard resistor were recorded at intervals as close as 0.5 JLS with a digital recording oscilloscope, 2048 points were recorded on each sweep of the oscilloscope on each of two channels giving a total sweep time of about 1 ms for the fastest rate. These data provided the carbon film temperature rise and the corresponding heat flux as a function of time. A minicomputer system processed the data and stored the results on magnetic tape for a permanent record. Accuracy of Results The carbon thermometer was calibrated at room temperature and by nitrogen and helium vapor pressure measurements. The maximum uncertainty in temperature measurement owing to calibration and recording instrument error is estimated to be ± 1.5%. Each time the carbon film was cooled from room temperature to helium temperature, the resistance of the carbon film was recorded. This resistance was found to be 2.6% less than the calibration resistance for the worst case. This offset, determined before each run, was assumed to be constant over the entire temperature range (up to 90 K) and was included in the calculations. Because of the offset and calibration error the total uncertainty in ~ T is estimated to be of the order ±5% . The heat fl.ux uncertainty is of the order ±0.4% arising from instrumentation error, error in surface area measurement, and accuracy of the standard resistor used in the measurement circuit. 458 P. J. Glarratano and N. V. Frederick TIME . s 4 .5 4 .0 3.5 N NE E 3.0 <.> ~ -.. 3: x:::> ... ..... ...es: -' 2.5 3: x ""' :::> ..... 2.0 4 1.5 ..... ...es: :r; :r; 1.0 0.5 TIME. s Fig. 3. Sampies of time-rlependent heat flux and temperature-difference data. PRESENTATION OF RESULTS Typical plots of controlled a T vs. time and the resulting q vs. time are shown in Fig. 3 for two different values of dT/ dt. Since a finite amount of power must be dissipated in the sample for the temperature controller to function, it is not possible to start control at a T = 0. As seen in Fig. 3 a T was generally of the order 0.6 K at t = 0. Consequently, the mode of heat transfer at the outsetwas already nucleate boiling. As the rate of change of temperature difference was increased, the various modes of pool-boiling heat transfer were observed. That is, the heat flux increased up to qmax (nucleate-boiling mode), decreased to a minimum, qmin (transition-boiling mode), and finally increased again beyond qmin (film-boiling mode). It was also determined that for rates of dT/ dt less than about 600 K/s the film-boiling regime was characterized by noise in the heat flux. This is illustrated in Fig. 3 for the slowest rate of heating. Figure 4 is a plot of heat flux vs. temperature difference for values of dT/ dt varying from 7500 K/s down to 0, i.e., steady state. As can be seen, the steady-state boiling curve bad a positive slope for all values of a T, i.e., even during the transition to film boiling the heat flux continued to increase. However, as dT/ dtwas increased, the negative slope portion of the q vs. a T curve began to emerge, becoming quite pronounced at dT/ dt = 7500 K/s. Under the transient condition, the peak nucleateboiling heat flux, qmax• the minimum film-boiling heat flux , qm;n, and the aT's at which they occur all increased with increasing dT/ dt. The most unexpected aspect of the data shown in Fig. 4 is the shape of the steady-state curve. The explanation for the absence of a negative-slope portion of the curve is not immediately obvious to the authors. It is planned to investigate heat transfer characteristics for surfaces other than carbon film and larger size samples to determine whether size or surface effect may account for this portion of the curve. It was noted, however, that for the steady-state data, there was virtually no evidence of Transient Pool Boiling of LHe Using a Temperature-Controlled Heater Surface 459 / 1.0 I 11 I Kutateladze ....._ ......_ Correlation P=0.82 atm I I /: I ,/0.[ I / " /1.: 0.1 I I I I I I (Zuber Correlationl · ···---- 7500 K/s 4000 K/s - ··- - · 2000 K/ s •··•· ···•· · 1000 K/s - - - 400 K/s - ····- 200 K/s - - 4 0 K/ s ar. K Fig. 4. Helium-boiling curve for various rates of heating. hysteresis in heat flux during several heating and cooling cycles. This was true under temperature control and heat flux control. As dTI dt was increased, hysteresis became evident and increased with increasing dT/ dt. This will be discussed in more detaillater in the paper. For comparison, Fig. 4 shows the nucleate pool-boiling correlations of Kutateladze [8 ] , the film-boiling correlation of Breen and Westwater [9 ], and the minimumfilm boiling heat flux calculated from Zuber's C0 J correlation. The deviation of these data from Kutateladze's predicted values is not unusual if one compares the other steady-state nucleate-boiling data in the literature with the same correlation, e.g., Brentari's [11 ] comparison of helium pool-boiling data with the Kutateladze correlation. The deviation of these data from the Breen and Westwater correlation for film boiling is more substantial, but not surprising, in view of the absence of a transition boiling regime for these steady-state data. At high fiT for the steady-state condition, the data asymptotically approach the Breen and Westwater predicted film-boiling values. DISCUSSION The concept of the existence of a thin macrolayer of liquid beneath an agglomeration of vapor bubbles during nucleate boiling near the peak heat flux has been suggested by many investigators. An excellent discussion of this view and a summary of the literature reporting the existence of the macrolayer is given by Yu P. J. Giarratano and N. V. Frederick 460 and Mesler [12]. Based on these observations, it has been assumed that a liquid layer of thickness 81 has been formed prior to the transient. A further assumption that has been made is that the thermodynamic and transport properties of the layer can be evaluated at the temperature of the heater surface, and that the time-dependent heat input, q(t), is partially dissipated into the bulk liquid above the macrolayer at a rate equal to the steady-state value, q••• corresponding to liT(t). The remaining heat input goes toward heating and vaporizing the liquid film, i.e., (3) where the first term on the right-hand side of (3) represents the sensible heat accumulated in the liquid layer while the second term involves the heat absorbed in vaporization of the liquid layer. As· can be seen from Fig. 4, this model is in qualitative agreement with the experimental data obtained in this study in that the transient heat transfer is greater than the steady-state value for a given 11 T, and the difference increases with increased rate of heating. It was concluded from the data that the total excess energy, I (q,- q••) dt, up to the time where minimum film boiling was established, tqmtn' was approximately a constant for all rates of heating (see Table I). This result lends further support to the concept of the liquid macrolayer. The initial thickness, 80 , was estimated by ignoring the sensible heat term of (3) and these values have an average value of 4 x 10-4 cm. Yu and Mesler C2 ] have reported thicknesses of the liquid macrolayer ranging from 4.8 x 10-4 cm to 9 x 10-4 cm for their nucleate-boiling data for water near the peak heat ftux. Since I (q 1 - q..) dt was a constant, as the time scale for the transient decreased with increased rate of heating, it follows that the difference between the transient and steady-state heat transfer must increase. This was observed experimentally. In spite of the highly simplified model described above, which does not take into account convective effects, it does appear to qualitatively predict theseexperimental results with the heat absorbed by the liquid layer accounting for the enhanced heat transfer under the transient condition. Iwasa and Apgar [4 ] found that their transient helium film-boiling data could be described by an expression similar to (3). The difference between their transient and Table I. Calculation of 8o Heating rate, K/s 7500 4000 2000 1000 400 200 100 40 J[q(t)- q,.] dt, J/cm 2 13.3 X 10-4 13.9 x 10-4 12.8 x 10-4 14.6 X 10-4 11.3 X 10-4 11.1 X 10-4 7.0 X 10-4 10.5 x 10-4 4.8 X 10-4 5.0 X 10-4 4.6 X 10-4 5.3 X 10-4 4.1 x 10-4 4.0 X 10-4 2.5 X 10-4 3.8 X 10-4 Avg = 11.8 x 10-4 Avg = 4.3 x 10-4 Transient Pool Boiling of LHe Using a Temperature-Controlled Heater Surface 461 steady-state heat transfer data was accounted for by the heat absorbed (on heating) by the vapor layer insulating the heater surface and the latent heat absorbed in vaporizing the bulk liquid above the vapor layer. This is consistent with the experimental transient film-boiling data presented here as is also evident from Fig. 4 where q(t) > q•• in the film-boiling region and where the difference q(t)- q•• increases with increasing heating rate. COMPARISON WITH DATA OF STEWARD An interesting comparison of this data during the heating cycle is made with Steward's [5 ] data for helium pool boiling in Fig. 5. First, however, it must be noted that Steward's data were obtained with a carbon film sample similar to the one used in this study but under different transient conditions. His data were obtained for a step input in heat ftux, and thus he recorded and presented temperature rise as a function of time for various constant heat ftuxes. By plotting the linear AT vs. time variations used in this experiment on Steward's plot and selecting the q's and AT's along this line, the q vs. AT plots for various dT/ dt could be obtained for Steward's data. The maximum upper bound on the error in q and AT owing to this plotting procedure is approximately ±50% for both variables. However, since a fairly smooth curve could be drawn through the data points so obtained, a more realistic estimate of the error is ±25% for both variables. There is a distinct similarity between the two sets of data which were obtained under different transient conditions. The similarity suggests that the heat transfer process is independent of the temperature history if the time to achieve the AT is the same. This matter requires moreextensive experimental and analytical investigation, but if substantiated, then a singleexperimental investigation using, e.g., a step in heat ftux as the transient condition could be used to predict q vs. AT under transient conditions where the temperature difference is the time-dependent variable. The similarity between the cross plot of Steward's data (obtained during a step in heat ftux) and the present data (obtained during temperature control) are also interpreted as evidence in support of the existence of the liquid macrolayer. That is, the energy required to vaporize the liquid layer (resulting in the transition to film boiling) is a constant and is independent of how the energy is applied. Analogaus to the model described earlier for a temperature-controlled transient, we may write for a heat-ftux-controlled experiment the heat balance as (4) That is, for heat ftuxes above the steady-state peak nucleate-boiling heat ftux, q•• m•x (below this heat ftux there would be no transition to film boiling) we assume again that the heat input, qin, is partially dissipated into the bulk fluid at a rate equal to q••max' and the excess heat goes to heat and vaporize the liquid macrolayer resulting in a transition to film boiling. Minimum film boiling occurs when sufficient excess energy has been transferred to completely vaporize the liquid macrolayer. The energy balance for the time to reach minimum film boiling is given by (5) where Oexcess represents the energy required to heat and vaporize the liquid macrolayer. Using Steward's experimental data for heat ftuxes which exhibited transition to film boiling [q ~ 1 W /cm 2 ], it was determined that qssmax = 0.95 W /cm 2 .,;. --3:... NE 3: .,;. ~ NE I ' I"' Stewa rd 's dcua, CIOSS plot of Obtained lrom I 0.011 01 I 1 I !I I I0 II I I K 1S I !! II LH . K 1000 I "I 10 111 I # I I t I I I I I I 1! I! I 100 ! 1 rl l /~ ··- <:T --3:... NE .,;. ... --3: NE 0.011 0.1 10 10 o o1 ! 0 1f 10 t · ~· I I I I I I I ' td °" I 1 11! Obtamed hom Cr OSS pl ol of Steward's dlttil I I I . . .. .. I I I I l I I I '" "I I I I I J !! . . _ Exper imental data. this study . .... I I I __/ llT. K Fig. 5. Comparison of helium-boiling curves obtained under two different transient heating conditions. ! ! /L 0.011 ::"IC Obtained IIom 1:·.~:.:~: 10 I '" J !!I I! I I II ::J j . . .. .. I "10:" :::!. ~ Cl. ~ '"'l ~ . ?'! Cl. "'= Q = ! ::: ;- ~ ~ ~ I ~ "" .". Transient Pool Boiling of LHe Using a Temperature-Controlled Heater Surface 463 and Oexcess = 14 X 10-4 J/cm 2 • Oexcess determined from Steward's data thus is in agreement with the values given in Table I for the temperature-controlled experiment. HYSTERESIS A limited series of experimental runs are available in which the hysteresis in heat ftux during heating and cooling cycles was investigated. Hysteresis, is defined here as the difference between q(t)heatinj and q(t)cooling at the same 6.T. Experimentally, attempts were made to maintain dT/ dtlheating = ldT/ dtlcooling and for some runs the Rare • 1490 Kl s Dweil Ralo · 2900 K/ s -- 0 Ms - "'E 150 Ms 10° <> ;: qs1udv s1a1e 1' x ........ ... I ::1 I I 1<[ : ... ' .... I 10 ' I I I <[ : I I I I I I, I I I - - ... J 10 -2 0.2 10 10 LH . K IH. K - 290 K/s N .... E -..<> E 3: <> ~ x ........ ... x ::1 ...... ..... ::1 1<[ .... : <[ : IH. K LH . K Fig. 6. Hysteresis data including the film-boiling region. P. J. Giarrataoo aod N. V. Frederick ...E <> ~ >< ... ..... ::I ... I- et ::z:: IH. K Fig. 7. Hysteresis data in the nucleate-boiling region. dwell time between heating and cooling cycles was varied from 0 to a time period sufficient to reach steady state. As noted earlier, the reported steady-state data exhibited no hysteresis. As the rate of heating increased, hysteresis began to emerge and increased with increasing heating rate. The degree of hysteresis also increased with dwell time between heating and cooling. The graphs shown in Fig. 6 show the experimental data exhibiting these trends. The temperature excursion was up to approximately 13 K, weil into the film-boiling regime for all values of dT/ dt noted. However, as shown in Fig. 7, for a temperature excursion in the nucleate-boiling region, little or no hysteresis was detected even at the relatively high rate of heating, 1780 K/s. The observed hysteresis in Fig. 6 is qualitatively understood in the context of the simple models discussed earlier. That is, q(t)- q.. = dT/dt > 0 on heating and q(t) - q.. = dT/ dt < 0 on cooling. This suggests that the degree of hysteresis for the fast transients is greater than that for the slower transients. At dT/ dt = 0 (steady state) the models predict q(t) = q•• for both heating and cooling and thus agrees with the present experimental observation of no hysteresis for the steady-state condition. The effect of dwell time between heating and cooling cycles, and the absence of hysteresis in the nucleate-boiling regime, even for a fast transient, Fig. 7, is not readily understood in terms of the simple models. A more detailed experimental investigation and careful analysis of hysteresis during heating and cooling is required. SUMMARY OF RESULTS The entire boiling curve, q vs. ~ T, was obtained for various values of dT/ dt. This was made possible through the use of an electronic temperature controller which was used to linearly vary in time the surface temperature of the carbon film heat transfer surface. Under steady-state conditions (dT/ dt = 0) there was no evidence of transition boiling in the usual sense, i.e., a decrease in heat flux with increased temperature difference. Surface effects or sample size may explain this result; however, further experimental investigation is required. Transient Pool Boiling of LHe Using a Tempenture-ControUed Heater Sudace 465 As dT/ dt was increased, the transition-boiling regime began to emerge, the peak nucleate-boiling heat ftux, the minimum film-boiling heat ftux, and the !iT's at which they occurred all increased with increasing dT/ dt (see Fig. 4 ). A simple model shows that the difference between the transient and steady-state heat transfer for a given !1 T is accounted for by the excess heat required to vaporize a liquid macrolayer of thickness approximately 4 JLm adjacent to the heater surface. Calculations showed that the energy required to completely vaporize the macrolayer (complete establishment of film boiling) is a constant of approximately 12 x 10-4 J/cm 2 • This was shown to be true for the present data obtained with temperature as the controlled variable as well as Steward's [ 5 ] data with heat ftux as the controlled variable. Hysteresis in heat ftux during heating and cooling cycles which extend iQ.tO the film-boiling region increased with increased dT/ dt with no hysteresis for the steady-state condition. This experimental Observation was also qualitatively explained by the model which predicts q(t)- q•• -= dT/ dt (greater than zero on heating, less than zero on cooling, and equal to zero for steady state). No hysteresis was observed in the nucleate-boiling region even for the higher heating rates. The model does not completely account for this result. ACKNOWLEDGMENTS The funding for this work was granted by the Air Force Office of Seienlilie Research, Bolling Air Force Base, Washington, D.C. with W. Dunnill, Program Manager. The authors also gratefully acknowledge the contributions of V. D. Arp, M. C. Jones, and W. G. Steward through many helpful discussions regarding this work. NOTATION CP = specific heat at constant pressure, 1/g-K dT/dt= rate of change of temperature, K/s h = heat transfer coefficient, W /cm 2 -K k =thermal conductivity,.W/cm-K q = heat flux, W/cm 2 t =time, s ~ T = surface temperature minus bulk temperature, K Greek Symbols a = thermal diffusivity, cm 2 /s 8 = thickness, cm A = latent heat of vaporization, J I g p = density, g/cm 3 Subscripts f =fluid i = incip = in= I= max = min = ss = s = insulation incipient nucleate boiling input liquid maximum nucleate boiling minimum film boiling steady state substrate 466 P. J. Giarratmo and N. V. Freclerick REFERENCES 1. J. Jackson, Cryogenics 9:103 (1969). 2. 0. Tsukamoto and S. Kobayashi, J. Appl. Phys. 46:1359 (1975). 3. R. L. Bailey, "Heat Transfer to Liquid Helium in Pulsed Heated Channels," Experiment Report RL-73-089, Dept. of Engineering Science, Rutherford High Energy Laboratory, Chilton, Didcot, Berkshire, England (1973). 4. Y. lwasa and B. A. Apgar, Cryogenics 18:267 (1978). 5. W. G. Steward, Intern. J. Heat Mass Transfer 21:863 (1978). 6. C. Schmidt, Appl. Phys. Lett. 32:827 (1978). 7. L. C. Brodie, D. N. Sinha, J. S. Semura, and C. E. Sanford, J. Appl. Phys. 48:2882 (1977). 8. S. S. Kutateladze, Fundamentals of Heat Transfer, Academic Press, Inc., New York (1963). 9. B. P. Breen and J. W. Westwater, Chem. Eng. Progr. 58(7):67 (1962). 10. N. Zuber, M. Tribus, and J. W. Westwater, "Hydrodynamic Aspects of Boiling Heat Transfer," AECU-4439 (1959). 11. E. G. Brentari, P. J. Giarratano, and R. V. Smith, NBS Tech. Note No. 317 (1965). 12. C.-L. Yu and R. B. Mesler, Intern. J. Heat Mass Transfer 20:827 (1977). DISCUSSION Question by J. W. Westwater, University of Illinois: The boiling heat flux at any ~ T is determined in part by the fluid flow pattern set up at that ~ T. Most of the runs described in this paper were completed in less than one second. This seems too short a time to allow for the establishment of any flow pattern characteristic such as the peak heat flux. What rate of changing the heater temperatures do you think will give an unsteady-state boiling curve which is equivalent to the steady-state boiling curve? Answer by author: All of our experimental data show that for heating rates less than 40 K/s, the unsteady-state boiling curve is essentially equivalent to the steady-state boiling curve. Comment by P. Seyfert, Centre d'Etudes Nucleaires de Saclay, France: The model suggested in this paper implies a macrolayer of fluid ad jacent to the surface with a heat flow independent of initial thickness. The Observations presented by Tsukamoto et al. * suggest, on the other hand, a liquid layer with initial thickness strongly dependent on heat flux. Even though, different fluids were investigated (helium and nitrogen) could you comment on this contradiction? Answer by author: We cannot presume to explain Tsukamoto's observed heat flux dependence of the vapor film thickness which was formed at the onset of film boiling in his experiment. However, if we assume that it is reasonable to compare their vapor film thickness with the nucleate-boiling liquid film thickness of our study, then we pointout that the mass equivalent of liquid thickness which would produce the vapor film thickness reported by Tsukamoto ranges from 1.4 x 10-4 to 2.9 x 10-4 cm as given by [Bvapor(Tsukamoto) X (Psatliquid/Pvapor@ Tttansition)]. This is the same order of liquid thickness calculated from the modeland helium data of our paper, namely, 4.3 x 10-4 cm. Question by D. Petrac, Jet Propulsion Laboratory: A carbon film is used both as a heater and as a thermometer in your measurements. Did you consider the effect of reflections from the carbon layer in the determination of the ~T since the calibration was performed at different power inputs (current Ievels) than during the measurements? Answer by author: Since the presentation of this paper, we have conducted tests on carbon film samples similar to the one used in this experiment to determine more conclusively the current dependence of the resistance. This was achieved by installing an independent heat source between the lower surface of the substrate and the insulation. The carbon film covered the entire top surface of the substrate. For various values of power input to the independent heat source, the resistance of the carbon film was recorded using a 2 x 10-6 -A current source. Then, with the independent heat source turned off, the power input to the carbon film was varied and the corresponding resistance recorded. A plot of resistance vs. power Ievel for the two cases was the same. We conclude that for these films the resistance is independent of the current Ievel through the film and depends only on the temperature. * 0. Tsukamoto, K. Uyemura, and T. Uyemura, in Advances in Cryogenic Engineering, Press, New York (1980), p. 476. Vol. 25, Plenum H-5 HEAT TRANSFER DURING SUBCOOLED HYDROGEN BOILING B. I. Verkin, Yu. A. Kirichenko, and N. M. Levchenko Academy of Seiences of the Ukrainian SSR Kharkov, USSR INTRODUCTION Various applications of cryogenic liquids in engineering have created continuous interest in the heat transfer processes occurring during subcooled boiling. Experimental studies of pool boiling of subcooled cryogenic liquids are few in The number. These include studies with hydrogen CJ. nitrogen [2 ], and helium available data are insufficient to determine the subcooling effect upon the heat transfer coefficients. The limited information which is available on the vapor bubble dynamics in boiling nitrogen was obtained with very low subcooling, i.e., about 1 K [ 4 ]. In this work the vapor bubble dynamics and heat transfer during subcooled hydrogen boiling were studied experimentally in the pressure range from 0.072 to 5 bar. A model is proposed to explain the heat transfer mechanism operating in subcooled boiling. Based on the model and experimental data, a relationship has been derived which adequately describes the heat transfer in boiling hydrogen with maximum subcooling for p:::; 3 bar. eJ. EXPERIMENTAL CONDITIONS Bubble growth and detachment were studied during subcooled hydrogen boiling with the aid of high-speed photography at pressures ranging from 0.072 to 1 bar, and subcooling (Tsat- T1) ranging from 0 to 6.4 K. Heating was provided by a dc current passing through the steel and copper-nickel horizontal tubes, 3 to 5 mm in diameter. The temperature of the heater wall and the liquid was measured with platinum resistance thermometers. Subcooling was provided by helium supercharging. The mean bulk temperature of the liquid was maintained constant and equal to the triple point by keeping it in contact with solid hydrogen. Thus, the experiments were carried out under maximum permissible subcooling for each pressure. GROWTH AND DESTRUCTION OF V APOR BUBBLES The growth and destruction of vapor bubbles in subcooled boiling hydrogen are These results are shown in Fig. 1, where described by Daugherty-Rubin's theory the relative subcooling and reduced pressures are CptJ/ L < 0.08 and pj Pc < 0.03, eJ. 467 468 8.1. Verkin, Yu. A. Klrichenko, and N. M. Levchenko respectively. The equation for bubble growth is given by R = ß-r" with n = 0.5, and the growth modulus is ß = cß"a n I ßya = cß (1) (A-~Tsat)"ß ,-va Lpva (2) with n13 = 0.75 and C13 :::::: 2. Observations made of hydrogen boiling under large subcooling conditions (I} > 4.5 K), showed that individual bubbles became attached to the heater and their size remained constant during the entire period of filming (0.2 s). DETACHMENT CHARACTERISTICS The pressure dependence of the vapor bubble detachment radii Rd is shown in Fig. 2 along with data for saturated hydrogen boiling [6 ' 7 ]. For subcooled boiling at 0.072 to 0.2 bar Rd.ub :=:::: RtJ..,; but for p > 0.2 bar Rd.ub < RtJ.., (the indices "sub" and "sat" refer to subcooling and saturation, respectively). The data obtained in this study may be qualitatively (and for Rd even quantitatively) interpreted in terms of the dynamic and quasistatic conditions of vapor bubble growth and detachment [8 ]. At low pressures the detachment characteristics t (( M• 0 ·1 !; (0 Ql • ·z D-j ~ 0~ t-.._ 0 \ llI !lt.o.<_ (0 tl'! ll6 ~ " \. _, lo 1-- ).a, ...... 1-o--, 1'---P.k. 9 6 Fig. 2. Departure radius of vapor bubble vs. pressure for boiling hydrogen. Saturated liquid: 1 and 2, present results; 3, Bewiloqua et al. ['], subcooled boiling. 469 Heat Transfer during Subcooled Hydrogen Boiling a8 -* • - z 00 _, 6 - " (Ö) - 6 Q- 1 ~ 0 -~ Fig. 3. Tmax/ Tc vs. reduced subcooling Cp t~/ L. Legend : 1, water, reference 10; 2, reference 10 (p = 1 bar) ; 3, hydrogen, present results; 4, nitrogen, present results ; 5, reference 9 (p = 1 bar); 6, unexplained results. T / /~ _._ 1\ _T_ . . .J-• '? ~~ 0 ~ L (}()} {]Oll Q(Jd tXM (J.f tll2 are governed by dynamic forces , i.e., forces of the liquid inertial reaction [8 ] . Thus, _ C2ß2; 3g-2; 3 R d _- c R ß4! 3g -1 / 3, (3) 'Td T At higher pressures the ftow changes to the quasistatic regime [8 ] and the Rd value can be found from the buoyancy-surface-tension equilibrium relation given by [3 J1/ 3 - R cU R _ _ d 2 g(pi - Pv) _ [ 3bU 2T.sat ]1/ 3 g(pt - Pv)flTsat ' (4) where Re = bR* is the radius of the microcavity from which the bubble departs, and R* is the critical radius of the vapor nucleus. Experimental data for Rd and 'Td with saturated and subcooled hydrogen boiling (Cp-&1 L < 0.08) are generalized by (3) and (4) for CR : : : : 1, bsat::::::: 15, bsub::::::: 4, and the growth moduli calculated by (2). Maximum errors are estimated tobe not more than ±30%. Daugherty-Rubin's theory [5 ] is helpful in describing the time dependence of the relative magnitudes of bubble radius (RI Rmax vs. Tl Tc, where Tc is the time covering the period from bubble nucleation till its complete destruction), if the parameter 'Tmaxi'Tc is known. In Fig. 3 'Tmaxi'Tc is shown vs. the relative subcooling Cp-&I L along with data for boiling water [9 ' 10 ]. 'T maxi 'Tc is seen to vary from 0 to 0.5 as CptJIL approaches 0.1. Note that 'Tmaxi'Tc = 0.5 is the limiting value in DaughertyThe curve in Fig. 3 can be employed to obtain a rough theoretical Rubin's theory estimate of the relative magnitude of bubble evolution. Absolutevalues of R(T) can be derived if some technique for evaluating Rmax and 'Tmax is available. For the entire range of the parameters studied Rmax ::::::: Rd; the error is estimated tobe no more than ±5% . The relation between 'Tmax and 'Td is shown in Fig. 4, where the data for water arealso given. 'Tdi'Tmax increases for greater subcoolings. Once 'Td is known, 'Tmax can be calculated from relation eJ. (5) Fig. 4. Td/Tmax vs. reduced subcooling Cp{}/ L. Water: 1, reference 10 (p = var); 2, reference 10 (p = 1 bar). Hydrogen: 3, present results. Nitrogen: 4, present results. aoz 470 8.1. Verkln, Yu. A. Klrlchenko, and N. M. Levchenko BEAT TRANSFER MODEL To develop a heat transfer model assume that the heat flux density q during subcooled nucleate boiling consists of three components, (6) where q 1 is a fraction of the heat flux density consumed by vaporization atthe heater, q 2 is the fraction expended on heating the cold liquid replacing the bubble that has just departed (or condensed) from the heater, and q3 is the fraction of the total heat flux density removed through natural convection. The latter is negligible for developed nucleate boiling. We can write (7) and (8) Here, Ii T 1 is the temperature gain in the "cold liquid" replacing the departed bubble. In the first approximation liT1 = Tw -11 2 Tt = O.S(Tw - Tt) = O.S(IiTsat + ~) On substituting (7), (8), (9) into (6), provided that q3 qsub (9) = 0, one obtains = ~7TR~fLpvz(1 + O.S.Ja + O.S.fasub) (10) where di _ "'a- Pt Cp Ii Tsat L Pv ' di _ PtCp~ eTasub- L Pv Within the model considered, (10) suggests that during saturated liquid boiling (~ = 0, .faaub = 0) and at sufficiently high pressures (.Ja < 2) the heat transferred from the heater to the liquid is mainly consumed by liquid evaporation at the heater {q1 > q 2 ). During saturated liquid boiling at low pressures (.Ja > 2) and boiling with considerable subcooling (.Ja + .Jasub > 2) the energy from the heater is expended on heatingthe liquid {q2 > q 1 ). This mechanism of "exchange" between the "bot" liquid of the boundary layer at T1(Tw > T1 > 11) and the "cold" liquid of the core at Tt accounts, as will be shown below, for the pressure dependence of the heat transfer coefficients arising during subcooled boiling. The model developed permits the heat flux fraction for evaporationtobe found. From (7) and (10) it follows that qsub 2 + .Ja + cfasub (11) For ,Ja + .fasub » 1 (Iow pressures and considerable subcooling at both high and low pressures) (11a) Heat Transfer during Subcooled Hydrogen Boiling 471 As seen from (11) and (11a), the heat ftux fraction for vaporization increases as pressure rises and the total temperature drop (Tw - T 1) decreases. The model proposed is employed to relate the heat transfer coefficients for saturated and subcooled liquid boiling. Equation (10) leads to the relation between the heat ftux densities for subcooled and saturated liquid boiling qsub = R~;ubfsubZsub(2 +Ja + Jasub) 2 +Ja R d ••.fsatZsat qsat For similar temperature drops (d Tsub (12) = d Tsat) (13) The expression in the brackets is readily calculated for a given d Tsat· This can be simplified for the limiting cases of very high or very low pressures as PvCpiJ 1+ -2Lpv (14) for Ja « 1 and {} 1+- (15) aT, for Ja » 1. To evaluate the factor in front of the brackets in (13) requires a knowledge of the boiling molecular characteristics of saturated and subcooled liquids (Rd, ß, Z). If the time dependence of the bubble radius obeys (1), the growth modulus is represented by (2) [ 6 ], the frequency f being 1I T d, and the density of the vaporization centers Z is found from (16) where Cz is a constant about 10-7 to 10-8 and nz becomes = 2 [11 ]. Equation (13) then (17) = For similar q (qsub qsat), asub/ asat can be estimated to a first approximation, if the assumption of q- aT;at is made. Then ( CXsub) asat q = (CXsub) 113 asat (18) .6.Tsat Weshall calculate asub/ asat for boiling hydrogen using (13), (17), and published data on boiling molecular characteristics. According to Judd and Merte C2 ], the density of vaporization centers seems to be weakly dependent on subcooling, and hence for all cases it is assumed that Zsub = Zsat· Substituting (2) and ßsat = 4Ja 0 · 5 ~ [ 6 ] into (17) and taking into account that (19) 471 2 ~7 tO 8 6 64ft)'' - - ~ :; * 2 8. I. Verldn, Yu. A. Kiricbenko, and N. M. Levcbenko 2 -.Q._ Fig. 5. Dependence a 1ub/a .., on pressure. Experiment: 1, = 104 W/m 2 ; 2, q = 2 x 104 to 6 x 104 W/m2 • Calculation: 3, equation (21). 0 * 6 8f0' P.la, * 2 q b8 which follows from Fig. 2 at p > 0.4 bar, one obtains = (asub) (q•ub) qsat 4T.., asat = 0 _125"al/2(t + 4T.., "a "asub) 2+ (20) Having taken equation (17) into account, the relation for the heat transfer coefficients at constant q is now given by fh ( asub) asat q "a = 0.5"al/6(1 +~) 2+ 1/3 (21) Figure 5 plots equation (21) and presents appropriate mean experimental data. To make the calculation more convincing we assume "a = 2"ao, where "ao is calculated from the temperature drop corresponding to the onset of liquid boiling. Figure 5 shows that the model that was developed adequately describes the experimental data up to a pressure of 4 bar; the maximum deviation of the calculated dependence, curve 3, from the experimental points is about ±20%. At a pressure of 5 bar the experimental and calculated values differ by a factor of 2. The estimates obtained (Fig. 5) permit the conclusion that the model provides a reasonable qualitative description of the relation between the heat transfer coefficients for boiling subcooled and saturated liquids. In fact, the calculation using (17) and (21) together with the experimental data give similar qualitative results: aaub = aaat for p < 0.3 bar, asub < asat for p > 0.5 bar. The relation decreases with increasing p. The quantitative agreement between the experimental data and the model to within ±20% at p = 0.072 to 4 bar can be considered satisfactory for such a rough estimate. Experimental data showing the pressure dependence of the heat transfer coefficients aaat and asub are plotted in Fig. 6, where the heat transfer coefficient variations are presented in relative magnitudes a (p )/ a (p = 0.5 bar). Normalization to a (p = 0.5 bar) was chosen because at p < 0.5 bar, asat = asub as the experimental results suggest. As seen in Fig. 6, the difference between the heat transfer coefficients aaat and aaub is quite pronounced under high pressures. At p = 5 bar they differ by a factor of 2 to 4. ,f (P, « (t~ I • -1 0 -2 ~ I(! 6 6 / V 5 V -j l> _, 2 ~ vt A .6 F1'" 'i. *6610' 0 V P.kr 2 4 6611l' 2 * b8 Fig. 6. Dependence of relative heat transfer coefficients on pressure in boilin} hydrogen. Subcooled boiling: 1, q = 1 x 104 ; 2, 2 x 10 ; 3, 4 x 104 ; 4, 6 x 104 W/m2 • Saturated hydrogen boiling: 5', q = 1 x 104 ; 5, (2 to 4) x 104 W/m2 ; 6, calculation by equation (23). 473 Heat Transfer during Suboooled Hydrogen Boillng HEAT TRANSFER COEFFICIENTS DURING HYDROGEN BOILING AT IDGH PRESSURES Tests reported above on the proposed model show that its mechanism of heat removal from the heater is sufficient to account roughly for the dependence of the heat transfer coefficient on pressure and subcooling. During subcooled hydrogen boiling the heat transfer coefficients can be found using (8). Using (9) and (16) with nz = 2, and/ = 1/Td = ß 2 /R~ into (8) Ieads to _~ C q - 37r R C2C A(Lpv) 11 \Cppl) 312 (äTsat + "') äT!:f6 (3 z 4/3T5/3 [{p ) ]1/3 U sat I Pv g (22) Assuming a slight difference between the temperature drops for saturation and subcooling and taking into account that asub = qsub/ ä Tsat. a less precise but more compact and convenient relationship is obtained after some transformations neglecting the magnitudes of low powers as asub A 1/3 ( cp/)1) 1/2( = B1[(PI _ Pv ) g ]1/9 -T. U sat äTsat +"' )1/3 2/3 (23) q Here B1 = B · F, where B = ~~7rCRC~Cz) 113 = 0.34 x 10-2 with Cz B = 1.7 x 10-2 with Cz = 10- , and F defined by F = (LPv äTsat) = 10-8 and 1118 (24) uTsat can be considered a dimensional constant owing to its weak dependence on pressure. Inserting the mean value of ä T sat for each pressure, F varies from 2.5 to 2.75 cm - 1118 in the range of pressures from 0.3 to 5 bar. Its maximum deviation from the mean value is ±5% . In the calculation using equation (13) äTsat was used to denote the temperature drop during the saturated liquid boiling. In this case an error of about 5% is permissible; the error results from factors F and (äT. + "'). A more accurate calculation of heat transfer may be made using (22). The experimental data on heat transfer in subcooled boiling hydrogen are generalized (Fig. 7) using (23) in terms of the coordinates asub/ A, q, where asub is the experimental heat transfer coefficient and A is the set of thermophysical properties in (23), i.e., A = asub/ Bq 213 • 2 lJ ~ R 6 0-J lf t;,. - ~ 2 Fig. 7. Generalized experimental results for subcooled hydrogen boiling, equation (23). Legend: 1, p = 0.072 bar; 2, p = 0.14 bar; 3, p = 0.25 bar; 4, p = 0.54 bar; 5, p = 1 bar; 6, p = 21bar; 7, p = 3 bar; 8, p = 4 bar; and 9, p = 5 bar. b'~ D- f 0-2 8 ~ V~ ~ /a ~~ ~ Vo' r:s ~ 2 .~ ~ 0-5 V -6 r-.~~"' ·y 6 ·oo D- ' 0 -8 ~IN' 0 -g 9 "IM' 10' 2 4 " 8 tJ' 2 4 "8 tJ' 2 474 8. I. Verkin, Yu. A. Kirlchenko, and N. M. Levchenko As seen in Fig. 7, for q > 7 x 103 W /m 2 the maximum error in the generalization is ±25% for the pressure range of 0.072 to 3 bar. The experimental coefficientB is about 0.5 x 10-2 • At p = 4 to 5 bar, the heat transfer coefficients depart from the generalized dependence. The difference between experiment and the calculation utilizing (23) for p > 3 bar is also illustrated in Fig. 6. At p :53 bar the calculated dependence essentially coincides with the mean experimental results. This is in agreement with the data in Fig. 5, which shows that for p > 4 bar the estimate by the model differs sharply from experimental results. The fair amount of discrepancy between experiment and theory at p > 3 bar suggests that either some additional aspects of the process other than evaporation in the bubble and "liquid exchange" between the boundary layer and the cold core should be considered under relatively high pressures and subcooling, or that the expression for the molecular characteristics of boiling at p :5 1 bar should not be extrapolated to the pressures greater than 3 bar. NOTATION A = Olsub/ Bl a = thermal diffusivity, m 2 /s b = constant CP = isobaric specific heat, J/kg-K eR. c,., Cz = constants g = gravitational acceleration, 9.81 m/s 2 L = latent vaporization heat, J /kg nß, nz = powers p = pressure, bar q = heat flux density, W /m 2 R = bubble radius, m Rmax = maximum bubble radius, m T = temperature, K ~ T = temperature drop, K Z = density of vaporization centers defined by equation (16) 13 clh Greek Symbols 01 = heat transfer coefficient, W /m 2 K ß = growth modulus A =thermal conductivit~, W /m K p = mass density, kg/m fJ = subcooling, K r =time, s T max = time Of maximum growth, S u = surface tension coefficient, N/m Subseripts C = critical value d = detachment I= sat = sub = v = w= liquid saturation subcooled boiling vapor heater wall REFERENCES 1. C. F. Sindt, in Advances in Cryogenic Engineering, Vol. 19, Plenum Press, New York (1974), p. 427. 2. B. Mazurek, R. Gruca, and J. Rutkowski, in Proc. 6th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1977), p. 230. Beat Transfer during Subc:ooled Hydrogen Boning 475 3. H. Tsuruga and K. Endoh, in Proc. 5th Intern. Cryogenic Engineering Conference, IPC Science & Technology Press, Guildford, England (1975), p. 262. 4. H. C. Hewitt and J. D. Parker, Trudy AOIM Teploperedacha 90c::22 (1968). 5. D. E. Daugherty and H. H. Rubin, in Proc. 1963 Heat Transfer and Fluid Mechanics lnst., Stanford University Press, Stanford, California (1964), p. 222. 6. Yu. A. Kirichenko and N. M. Levchenko, Zh. Prikl. Mekh. Tekh. Fiz. 1976(4):103 (1976). 7. L. Bewiloqua, W. Görner, R. Knöner, and H. Vinzelberg, Cryogenics 14:516 (1974). 8. Yu. A. Kirichenko, Inzh. Fiz. Zh. 25(1):5 (1973). 9. L. Leppert and K. Pitts, in Prob/emy teploobmena, Atomizdat, Moscow (1967), p. 142. 10. D. M. Kostanchuk, thesis for candidate of Technical Seiences degree, Kiev, ITTF AN Ukr. SSR (1971). 11. K. A. Zhokhov, in Aerodinamika i teploobmen v rabochikh elementakh energooborudovania, TsKTI, Leningrad (1969), p. 131. 12. R. I. Judd and H. Merte Jr., Int. J. Heat Mass Transfer 15(5):1075 (1972). H-6 OBSERVATION OF RUBBLE FORMATION MECHANISM OF LIQUID NITROGEN SUBJECTED TO TRANSIENT HEATING 0. Tsukamoto and T. Uyemura Yokohama National University Yokohama, Japan and T. Uyemura Tokyo University Tokyo,Japan INTRODUCDON Several photographic observations of steady-state pool boiling of liquid nitrogen have been published C-2 ]. However, there are few papers reporting the Observation of transient-state boiling. Also, there is little information on bubble formation mechanism or heat transfer characteristics of the transient-state boiling in ~ a cryogenic liquid. This paper presents optical observation of bubble formation at the surface of a thin wire, heated by a stepwise transient in liquid nitrogen. Bubble formation was observed by a high-speed camera with a frame speed of 5000 frames/s. Interesting phenomena, peculiar to transient heating, were observed. Usually, it is assumed that when a wire is heated with a stepwise increase in the current, the boiling at the wire surface is nucleate for a short while after the step current has passed; eventually the boiling moves into film boiling, provided that the heating power per unit surface area of the wire is above the steady-state bornout heat flux. This follows because part of the Joule heating power is absorbed in changing the enthalpy of the wire at the beginning of the heating. However, it was observed that the boiling at the surface of a wire immersed in a saturated liquid nitrogen bath went directly into the film-boiling state without passing through the nucleate-boiling state. This phenomenon occurred even when the wire surface heat flux was below the steady-state bornout heat flux. Recently, Shinha et al. [3 ] reported on experimental results of transient heat transfer characteristics in liquid nitrogen, where they referred to a "premature" transition to film boiling. They verified that this is the case by measuring the heat transfer characteristics associated with the process. Highspeed photographs of such a premature transition to the film boiling have been made in this study. In addition, the transient heat transfer characteristics associated with 476 Obse"ation of Rubble Formation Meehanism of Liquid Nitrogen 477 the process were measured. This paper discusses the bubble formation mechanism for the transient heating and its relation to the heat transfer characteristics. EXPERIMENTAL ARRANGEMENT The experimental arrangement is shown in Fig. 1. A platinum wire was immersed horizontally in a pool of liquid nitrogen (at atmospheric pressure) contained in a 10-cm-ID glass dewar. The platinum wire was mounted under spring tension. The length and diameter of the heating wire were 23.5 mm and 0.05 mm, respectively. By applying a step voltage, a step current was fed to the wire through a current-limiting resistor. The resistance of the current-limiting resistor was considerably !arger than that of the wire so that the current could be regarded as constant in the range where the wire temperature was below 200 K. Even though the amount of Joule heating varied with the change in resistivity of the wire, the driving step current was sufficient for the present experiment. Current and voltage waveforms of the wire were recorded by digital transient recorders. The bulk temperature of the wire was estimated from the value of the resistance of the wire. The surface temperature of the wire was assumed to be equal to its bulk temperature because the wire was so thin that its cross-sectional quartz tlmer glass dewar current waveform T. R. 2 valtage waveform 500W mercury lamp hlgh-speed liquid N2 movle camera platinum wire C. R. current limiting resistor T. R. digital transient recorder Fig. 1. Schematic of experimental arrangement. 478 0. Tllllbmoto IUld T. UyeJHI'II temperature distribution could be regarded as being uniform. The transient heat ftux q(t) transferred from a unit of wire surface into the liquidnitrogenwas determined from the following equation: q(t) = __!__ [v(t)i(t) _ pc(T).".{~) 2 d7l 'TI'd L 2 dtJ (1) Differential values of the wire temperature were determined by a graphical differential method. The photographs were taken with a 16-mm high-speed movie camera. In the experiments, the movie camera was started 0.6 s before a step current was fed to the wire initiating the transient heating event. The framespeedwas about 5000 frames/s when the wire current began to build up. The starting times of the movie camera and the transient heating event were controlled by a quartz timer. The precise frame speed and event start time were determined from time marks on the film. Lighting was provided from the rear by a 500-W mercury lamp diffused through a frosted glass. The Ievel of liquid nitrogen was kept several centimeters above the platinum wire throughout the experiments. All data and photographs were taken at atmospheric pressure. RESULTS AND DISCUSSIONS The experimental steady-state bornout heat ftux was 17 W /cm 2 • For step heating, direct transitions to film boiling occurred whenever the surface heat ftux was more than 50% of the steady-state bornout heat ftux. A typical example of experimental current and voltage waveforms which were obtained during one such direct transition to film boiling is shown in Fig. 2. At t = 0, a step current was fed to the wire. The critical current icr above which the direct transition to film boiling occurred was 0.90 A and the corresponding Joule heating power per unit surface area of the wire was 8.2 W I cm2 for a temperature difference of 25 K. Temperature changes in the wire for various values of the step current are shown in Fig. 3. For the case in which direct transition to film boiling occurred, the temperature rise of the wire was small for a short time period after the step current bad been fed to the wire. Then, the wire temperature began to rise rapidly [curves (a), (b), and (c) in Fig. 3]. At a current near icr, the transition can be either sametime to film boiling or to nucleate boiling. For the case where the transition to nucleate boiling occurred, the wire temperature initially approached a steady-state convective value and then decreased when the nucleate boiling began [curve (d) in Fig. 3]. (Usually, the transition to nucleate boiling took more than several seconds after the initiation of heating, and the temperature trace when nucleate boiling beganisnot shown in the figure.) Fora smaller current, the wire temperature approached a steady-state convective value, and no boiling occurred [curve (e) in Fig. 3]. High-speed photographs of the liquid-vapor interface for the voltage and current waveforms of Fig. 2 are presented in Fig. 4. This is the case where a premature transition to film boiling occurred. For a short time after a step current bad been passed through the wire at t = 0, no change occurred around the wire. However, at t = t4 , a propagating vapor sheath quickly surrounded the wire [photographs (a) through (c) of Fig. 4]. The entire wire was surrounded with a vapor sheath in less than 3 ms after the first vapor appeared. After that, the surface of the vapor sheath began to undulate with the vapor sheath becoming divided into several Observation of Rubble Formation Mechanism of Liquid Nitrogen 479 2.0 > . ........,"' 11> ~ 1 .2 0 > 0 . ~ 0 50\ td 100 150 200 150 200 time , ms 1.0 <. .. 11> 11> 0. E "' 0.5 0 50 100 time , ms Fig. 2. Example of current and voltage waveforms during heating of wire. Joule heating per unit surface area of the wire is 10.4 W /cm 2 at t = td. parts [(d) through (f) in Fig. 4]. This process consumed about 5 ms. Each part of the divided sheath then grew to a bubble. The bubbles kept growing until they separated from the wire [(g) through (h) of Fig. 4]. Steady film boiling occurred thereafter. The bubble formation process, as shown in Fig. 4, was generally observed, when a premature transition to film boiling occurred. When the heat flux to the wire surface increased above the steady-state burnout heat flux, many small bubbles appeared almost simultaneously at t = td on the wire surface as shown in the photographs of Fig. 5. These small bubbles coalesced and a thin vapor sheath with undulated surface surrounded the wire. Then, the vapor sheath divided into larger bubbles and film boiling commenced. The time td when the vapor sheath surrounded the wire coincided with the time when the wire temperature began to rise abruptly. When the wire surface is surrounded by the vapor sheath, the heat is transferred through the vapor sheath and the heat transfer coefficient decreases abruptly. The value of td depends on the heating power. Curves of td vs. the heating power per unit surface area of the wire are shown in Fig. 6. Values of td varied considerably. Variations were larger as the heating power was decreased. Average thicknesses of the vapor sheath for various values of heating power per unit surface of the wire are listed in Table I. The mechanism of the vapor sheath formation is believed to be as follows: When a step current passes through the wire, liquid araund the wire is initially heated by the 0. Tsukamoto and T. Uyemun 480 (b) (c) 50 40 ., 00 "". 30 ~ (d) 20 (e) 10 100 50 150 200 t, ms Fig. 3. Temperature changes of the wire for various values of wire current. (a) I = 1.05 A (q = 10.5 W /cm 2); (b) I= 0.98 A (q = 9.6 W /cm 2 ) ; (c) I= 0.92 A (q = 8.5 W /cm 2 ); (d) I= 0.88 A (q = 7.2 W/cm 2 ); (e) I= 0.68 A (q = 4.1 W /cm 2 ) . (q in parenthesis is the value at t = td.) heat conducted from the surface of the wire CJ. Convection of liquid around the wire is not fully developed in a time as short as td. The liquid around the wire is superheated until, at td, it vaporizes abruptly. If this is true, the latent energy required to form the vapor e., which is contained in the vapor sheath, should agree with the energy ec which is transferred through the wire surface until the vapor sheath forms, provided that convection of the liquid is not completely developed (see Table I). ( a) t=O msec ( b) t= 54 . 6 msec (c) t= 55 . 8 ms e c . (e ) t=61. 3 msec \ f) t =6j.2 rn~ ec ( d ) t=59.4 msec - (g ) t =68 . o mc;ec (h) t= 73 . 8 ms e c Fig. 4. High-speed photographs of the vapor-liquid interface at the wire surface. Photographs show the case for the premature transition to film boiling (q = 10.4 W /cm 2 ). Observation of Rubble Fonnation Mechanism of Liquid Nitrogen - I (a) 481 t =O msec (b) t =4 . 6 msec t =4 .8 msec (c) (d) t=5.2 msec (h) t =29.0 msec • (e) t=7.0 msec (f) t=B. 6 msec t=16.6 msec (g) Fig. 5. High-speed photographs for higher Joule heating power (td = 4.6 ms}. Heating power per unit surface area of the wire is 35.1 W /cm 2 • 8oo 600 4oo 200 .. 100 80 U} 8 '0 .j..) ..: ·. 60 • ::· . ...,· 40 ·''t "'...1', 20 .. t ;. 10 8 I. \ "·\ 6 4 4 6 8 10 20 40 60 80 q , W/cm 2 Fig. 6. Time when vapor surrounds the wire, td, vs. the heating power per unit surface area of the wire. (The value of the heating power is the value at t = td.} 48l 0. Tsukamoto and T. Uyemura Table I. Energy for Formation of Vapor Sheath No. q*, W/cm 2 td,ms 68 v,mm 1 2 3 4 8.3 10.4 10.9 18.4 27.2 35.1 124 55.8 36.1 14.1 7.6 4.8 0.60 0.69 0.68 0.45 0.43 0.33 5 6 q* ec, 103 J/cm e., 103 J/cm 13.18 2.61 4.07 2.05 1.27 0.99 2.55 3.38 3.23 1.43 1.30 0.76 = Joule heating power per unit area of the wire. Except for the case where the time td was exceptionally long, the agreement between e, and ec is good. Preliminary experiments have been performed with liquid helium. Very little data are available at this time. Direct transitions or premature transitions to film boiling in liquid helium have so far not been observed. This is probably due to the much smaller surface tension of liquid helium where the latter detaches from the wire surface more easily than nitrogen bubbles. ACKNOWLEDGMENTS The authors gratefully acknowledge the technical advice of Y. Yamamoto, and T. Tsuno for taking high-speed photographs. The useful discussions with N. Inai are also appreciated. NOTATION C( T) d = specific heat of heater wire = diameter of wire e, = latent energy required to produce the vapor contained in a unit length of vapor sheath ec = energy dissipated into the liquid from the surface of the wire per unit length until the vapor sheath forms i(t) = wire current icr = critical current above which direct transition to film boiling occurs L = length of the wire q(t) = heat ftux dissipated into liquid per unit surface area of wire T = bulk or surface temperature of the wire .1 T = difterence between wire temperature and bulk temperature of liquid nitrogen td = time when the vapor sheath surrounds the wire or time when the voltage across the wire begins to rise abruptly v(t) = voltage across the wire Greek Symbols 6av = average thickness of the vapor sheath p = density of wire REFERENCES 1. R. J. Simoneau and K. J. Baumeister, in Advances in Cryogenic Engineering, Vol. 16, Plenum Press, New York (1971), p. 416. 2. E. R. F. Vinter, A. K. Vong, and P. MacFadden, J. Heat Mass Transfer 9:301 (1968). 3. D. N. Shinha, L. C. Brodie, J. S. Semura, and F. M. Young, Cryogenics 19(4):225 (1979), 4. P. J. Berenson, J. Heat Transfer 83, Series C (3):351 (1961). H-7 NATURAL CONVECTION HEAT LEAK IN SUPERCRITICAL HYDROGEN TANKS A. J. Barrett Beech Aircraft Corporation, Boulder Division Boulder, Colorado INTRODUCTION Information on natural convection, such as thermal pumping in large-diameter cryogenic lines, is weil documented C· 2 ]. However, less information is available on natural convection heat transfer in small-diameter tubes filled with supercritical fluid. Natural convection occurs in cryogenic systems where the warm tube interface (as shown in Fig. 1) is located below the cold tube tank connection. Natural convection heat leak increases the boil-aff rate; therefore, it is of practical importance to be able to predict its occurrence and to reduce the heat leak. This study was initiated to develop a method of predicting natural convection for variations in system geometry, fluid properties, and temperature gradients. The natural convection heat leak was determined from test data and an empirical correlation was obtained for the prediction of natural convection. DISCUSSION The natural convection considered here is similar to that described by Trucks C]. Consider a column filled with fluid at different temperatures, opening at the top into a tank filled with the same fluid at the colder temperature. The pressure at the bottarn of the column is unequal because of the density difference existing in the column. The pressure differential causes flow from the cold end to the warmer end. Furthermore, flow continues as long as the temperature differential is present to provide the driving force. Maintaining the temperature differential requires a continual supply of cold liquid from the tank and a heat input to the warm column. Steady-state recirculation flow rates and temperatures for natural convection systems are established by balancing system frictional pressure Iosses against the driving force caused by the density differential in the cohimn. This paper is concerned with the occurrence of natural convection in smalldiameter tubes that are filled with supercritical hydrogen. The warm end of the tube is at ambient temperature and is located below the cold reservoir. Free convection in vertical tubes has been studied C-4 ]. However, these past studies did not include supercritical cryogenic fluids. 483 484 A. J. Barrett flll TUBE-3/8" 0 D x .022 WALL X .022 WALL ~ GIRTH RING • IIY ..., /g. I j I ~ OINIIIY PRO I! Fig. 1. Upright tank configuration. TEST DATA Aseries of testswas run in which natural convection heat leak occurred. The test configuration consisted of a spherical hydrogen dewar with fill and vent lines going from the top of the pressure vessel to an equatorial girth ring (see Fig. 1). In addition to the upright configuration, the tank was turned upside down and the same tests were repeated. In the inverted condition, natural convection did not occur. TableI 485 Natural Convection Heat Leak in Supercritical Hydrogen Tanks Table I. Summary of Test Conditions* Test designation Ambient temperature, 0 R Cold fluid temperature, 0 R Cold fluid pressure, psia 1 2 3 4 5 6 7 8 9 10 570 570 570 570 570 570 515 515 515 515 48 66 48 66 48 66 48 66 48 66 280 280 280 278 280 283 257 262 250 260 Upright Inverted X X X X X X X X X X * Ambient pressurewas 12 psia. presents a summary of the test conditions. The two cold fluid temperatures are 48°R for the full condition and 66°R for the 34% full condition. Fluid pressures were maintained between 250 and 280 psia. A summary of heat leak values from the tests is shown in Table li. Total heat leak values were computed from the measured flow and the specific heat quantity which was determined from fluid temperature and pressure. Natural convection only occurs with the tank in the upright position, and the maximum value is 11.9 Btu/hr/line for a full tank. When the fluid quantity was depleted to the 34% full condition, natural convection heat leak was 3.3 Btu/hr/line. In order to determine the natural convection heat leak, it was necessary to predict the heat leak without convection and then subtract the prediction from actual test data. Heat leak predictions were obtained using an analytical hydrogen tank model that was correlated with the fluid condition of Test No. 10. The analytical modelwas made to Table II. Summary of Heat Leaks Test designation Baseline heat leak design, Btu/hr Natural convection heat leak, Btu/hr 4.9 23.8 (11.9 /line) 6.6 (3.3/line) 2 4.0 3 4 5 6 4.9 4.0 4.9 4.0 7 3.8 8 9 10 3.0 3.8 3.0 Pressure oscillation heat leak, Btu/hr Total, Btu/hr Tank quantity,% 28.7 100 10.6 34 100 34 100 34 23.8 6.6 80.3 4.4 149 4.0 109 15 7.9/line (10.8/line prediction) 4.2 2.4 22 144.1 2.8 2.2 10 6.2 3.0 100 34 100 34 l Ambient temperature = 570°R Ambient temperature = 515°R 486 A. J. Barrett agree exactly with Test No. 10. Also, by comparing test results (Nos. 7 and 9 for example), it is possible to obtain the natural convection heat leak. In a similar matter, it was possible to estimate the pressure oscillation heat leak in the tests. With a full tank, the oscillation caused heat leak was in excess of 144 Btu/hr, and as tank quantity was depleted, oscillation heat leak dropped below 5 Btu/hr. In the 515°R environment tests, the test configuration had a smaller diameter supply line outside the tank so that pressure oscillations were reduced. RESULTS Initially, heat leak data were correlated with a computer model similar tothat described by Trucks CJ. However, probably owing to the relatively low total heat leak encountered, satisfactory agreement was not obtained. Therefore, a dimensional analysisapproachwas used and the following expressionwas developed: . G = DCf[gpD 3 (Pc- Pa)(ha- hc)]N fik (Ta _ Tc} convectlOn (1) where an overbar denotes a value taken at arithmetic mean temperature. Using the above equation, a series of parametric curves was generated to show the sensitivity of 10° .; (.) c ü"' :> '0 c 10" 1 a: u ., 0 ~ 0 Cl c V ·~ ~ u-- G>:l > t- §co u (;; :; :;; z 10""' X T =560° R A T =660° R Cold fluid temp . T =48° R .50 .75 1.00 1.25 T ube diameter , inches Fig. 2. Effect of tube diameter on natural convection heat leak at 48°R. Natural Convection Heat Leak in Supercritical Hydrogen Tanks 487 10° .,·u c 10-' ~ u ;, "ca: u ., 0 "' 0c "~ ·z-'= u._ O>::J > 1- g!Il u ~ ::> ~ 10_, .50 .75 1.00 1.25 Tube diameter, inches Fig. 3. Effect of tube diameter on natural convection heat leak at 66°R. natural convection to tube diameter, temperature gradient, and fluid density. The results are shown in Fig. 2 where the natural convection conductance is plotted against tube diameter. This figure indicates that conductance varies as the fourth power of the tube diameter. The effect of ambient temperature seems tobe small. In Fig. 3, the same parameters are plotted. However, the cold fluid temperature was raised to 66°R, which lowers the fluid density. A comparison between Figs. 2 and 3 shows a reduction in heat transfer as the fluid density is reduced. The correlation between the empirical expression and test data is shown in Table II. By comparing the total heat leaks in Tests Nos. 7 and 9, a natural convection heat leak of 7.9 Btu/hr/line is obtained. The predicted value is 10.8 Btu/hr /line. In this comparison, it was assumed that the pressure oscillation heat leak was the same in each test. CONCLUSION An expression has been developed for the prediction of natural convection in terms of system geometry, fluid properties, and temperature environment. The correlation between test data and (1) is within 26% as shownon Table II, case 7. It would have been beneficial if more test data bad been available to substantiate the relationship. Equation (1) may be used during the design phase of supercritical fill 488 A. J. Barrett and vent lines to select the proper tube diameter to reduce the natural convection heat leak. NOTATION C = empirical constant (0.444 x 10-7 ) D = tube diameter, ft g = acceleration Ievel, ft/hr 2 h =fluid enthalpy, Btu/lb k = mean fluid thermal conductivity, Btu/hr-ft-•R N = empirical constant (0.984) T = fluid temperature, •R Greek Symbols p = mean fluid density, lb/fe ii = mean fluid viscosity, lb/ft-hr Subsaipts a = ambient c = cold fluid REFERENCES 1. H. F. Trucksand W. D. Randolph, in Advances in Cryogenic Engineering, Vol. 10, Plenum Press, New York (1965), p. 341. 2. S. K. Morgan and H. F. Brady, in Advances in Cryogenic Engineering, Vol. 7, Plenum Press, New York, (1962), p. 206. 3. D. W. Murphy, in Advances in Cryogenic Engineering, Vol. 10, Plenum Press, New York (1964), p. 353. 4. E. R. G. Eckert and T. W. Jackson, "Analytical Investigation of Flow and Heat Transfer in Coolant Passages of Free Convection Liquid Cooled Turbines," NASA TN 2207 (1950). DISCUSSION Question by B. A. Hands, University of Oxford, England: Did you examine the effect of a net mass flow through the tube? Answer by author: No. H-8 TECHNIQUES FOR REDUCING RADIATION HEAT TRANSFER BETWEEN 77 AND 4.2 K* E. M. W. Leung, R. W. Fast, H. L. Hart, and J. R. Heim Fermi National Accelerator Laboratory Batavia, Illinois INTRODUC'fiON The present-day applied superconductivity is a liquid-helium-based technology. The efficiency of a cryogenic or superconducting device is determined by its rate of consumption of the cryogen. Liquidhelium has an extremely small heat of vaporization; thus, its storage almost always requires a state-of-the-art insulating method. Radiation heat leak becomes significantly more important as larger superconducting devices are built, such as huge high-energy physics analysis magnets, fusion reactor systems, and energy Storage facilities. For large superconducting magnets, 3 ] it is common practice to surround the liquid helium vessel with a nitrogen shield (at 77 K) and wrap multilayer insulation around both the liquid helium vessel and the nitrogen shield in an attempt to further reduce the heat leaks (Fig. la). Multilayer insulation is inexpensive and generally effective; yet, it is a rather difficult material to apply because its performance depends on a few hard-to-control parameters such as the layer density, the compressive loading, and the lateral heat transfer effect. The effective insulation capability obtained in practice is at least a factor of 2 worse than carefully measured laboratory values (or those claimed by manufacturers). Careless and/ or inexperienced application can easily generate heat leak values a few times higher than the predicted value, especially when dealing with peculiarly shaped cryostats. Extensive study and test programs [4-6] have been carried out in an attempt to better understand the thermal performance of multilayer insulation but a Iack of experimental data still exists for temperatures below 60 K. In fact, the only experimental heat transfer data that appear to be available is a value of 14.3 mWym 2 (1.33 mW /fe) for a 60 layers/in. Linde system (Al foil + glass paper) [']. A commonly used design parameter taking penetrations and other imperfections into account is 43.0 mW /m 2 (4.0 mW /ft2) for NRC-2 insulation,t which is both the cheapest and easiest to apply of the multilayer insulations. In this paper, the 77- to 4.2-K performance of NRC-2 insulation as a function of the number of layers is presented. Results indicate that a much smaller number of layers is required to c- * Work sponsored by the U. S. Department of Energy. t Manufactured by King Seeley Company, Winchester, Massachusetts. 489 490 E. M. W. Leung, R. W. Fast, H. L. Hart, and J. R. Heim '".• """'" .",1I Wal ls of LN2 Vesse I ($ometimes Reploced by o Nitrogen Shield) Woll of the Vocuum Bo• Vocuum (a) L He Vessel wall ot 4 2K Hi<J> contoct resistonce, - - -...1 low thermal conductivity conn ecti ans (b) Fig. 1. (a) Typical arrangement of insulation system in !arge superconducting devices. (b) Floating-shield concept. obtain optimum performance, compared to the large number of layers (40 for minimum kp or 120 for minimum k) between 300 and 77 K. Pushing this concept of complete elimination of solid conduction to the Iimit would Iead directly to the classical calculation of Scott [8 ], who indicated an achievable heat transferrate of the order of 21.5 mW/m 2 (2 mW;te) between surfaces at 77 and 4.2 K, provided that both surfaces have sufficiently low emissivities (e = 0.01 to 0.03). Recent work of Chaussy et a/.[9 ] showed that a value of 18.7mW/m2 (1.74mW/fe) could be otained by covering both the 77-K and 4.2-K surfaces with a commercial aluminum adhesive tape, Eccoshield PST.* Substituting the emissivity values confirmed by Chaussy for aluminum (e77 = 0.03 and e4.2 = 0.011) into the parallel plate formula (for diffuse reftection) given by Scott [8 ] 6 A u(T~- T1) (1/ Et) + (1/ E2 - 1) (1) results in a heat transferrate of 16.5 mW /m 2 (1.53 mW /ft2 ). This paper describes a test program set up to further verify low radiation heat loss via this "reftective coating" method and to study the problern of degradation and contamination associated with such low emittance surfaces. The heat transfer from * Manufactured by Emerson Cuming, Inc., Canton, Massachusetts. Techniques for Reducing Radiation Heat Transfer between 77 and 4.2 K 491 77 to 4.2 K has been measured for cleaned copper to cleaned copper, Eccoshield Al tape to Eccoshield Al tape, 3M (Industrial tape # 425) Al tape to 3M Al tape. In addition, other ideas and experiments are also presented leading to a practical scheme for use in the Chicago Cyclotron Magnet (CCM), a large superconducting split solenoid, currently under construction at Fermilab. RADIATION HEAT TRANSFER AND EMISSIVITIES Success in reducing radiation heat transfer depends on achieving surfaces of low emissivities. Metals approximate gray bodies rather weil in the infrared region, consequently, one can take the spectral emissivity values tobe the same as the total emissivities without any appreciable error. The knowledge of low-temperature emissivity and absorptivity values is limited C0- 12]. Note that the emissivity value for a given material is also a function of the wavelength (Fig. 2) of the incoming thermal radiation (in other words, the temperature of the heat source). The emissivity of a practical surface may be quite different from the value predicted for the given material under ideal conditions since it is a strong function of chemical purity and surface roughness, contamination, and method of preparation. Engineering testing of practical reflective surfaces is therefore justified. Good electrical conductors such as silver, gold, copper, and aluminum have low infrared absorptivity. Gold is too expensive for practical use, while silver and copper tend to tarnish easily in air. Aluminum, with its natural protective layer of aluminum oxide (transparent to infrared radiation of temperature <100 K), is often used. Electropolished copper protected by a transparent benzotriozollayer, a few molecular layers thick, is also a good choice because the nitrogen shields in superconducting devices are typically made of copper whose high thermal conductivity helps to maintain a uniform temperature. The "floating-shield" concept is to insert a second reflective shield between the helium vessel surface and the nitrogen shield (see Fig. 1b). Theoretically, it can reduce the heat leak by a factor of 2. In practice, by attaching this secondary shield through high-contact resistance connections, one can reduce the heat leak by 35 to 40%. Radiation heat Ioad from 77 to 4.2 K of less than 10 mW /m 2 is achievable using this method. If one selects to cool this secondary shield with excess helium boil-off, one can easily reduce the temperature of the shield to around 50 K, in which case the heat Ioad to 4.2 K would be approximately 5 mW/m 2 • It has been suggested that a superconductor be used as a reflector. The emissivity should approach zero as the electrical resistivity vanishes. The Bardeen, ~ ~ Fig. 2. Dependence of the emittance of a surface on its temperature and the average wavelength of the incoming thermal radiation. 1.0 -~ Tz I T, ' - Temperoture of emilfing sur face Average wovelengtn of the incoming thermal rodiotion , ), AVG 492 E. M. W. Leung, R. W. Fast, H. L. Hart, and J. R. Heim Cooper, and Schrieffer (BCS) theory [u] suggests that foramaterial in the superconducting state there is an energy gap for electron states such that a threshold excitation is required before energy can be absorbed. This relationship takes the form of (2) Eg = hv = 3.5 kTc In other words, a superconductor of a critical temperature Tc would not absorb thermal radiation of the wavelength A > hc/3.5 kTc. The average wavelength of the thermal radiation from a surface at a temperature T is governed by the Wien displacement law (3) (AT)avg = 4107 ~m-K Simple arithmetic shows that Nb 3 Sn (Tc = 18.3 K, He = 1.85 x 104 kA/m), a superconductor considered having a rather high Tc by present-day standards, can only be a perfect reflector for thermal radiation from a surface of temperature less than 15 K. Hence, a superconductor is of limited value as an infrared reflector for most helium systems. TEST APPARATUS A drawing of the test setup is shown in Fig. 3. The 4.2-K surface A is a flat copper plate completely surrounded by a reetangular copper box B at 77 K. The sample to be tested is either attached (e.g., multilayer insulation), taped (e.g., Al 3.0 METE RS A SECTION A-A Fig. 3. Thermalradiationtest setup. Legend: A, 4.2-K fin; B, reetangular 77-K box ; C, inner helium vessel; D, outer helium vessel; E, nitrogen vessel; F, vent tube. Techniques for Reducing Radiation Heat Transfer between 77 and 4.2 K 493 tape), or coated onto both A and B. A capped-off 12.7-mm-ID copper tubesoldered to A is connected to the inner helium vessel C through a coupling. Any heat absorbed by the 4.2-K fin A rnanifests itself as liquid helium boil-off from C because all other sources of heat are intercepted by the outer helium vessel D and the radiation bafftes which almost enclose the inner helium vessel. The outermost concentric liquid nitrogen vessel E shields the outer helium vessel D from room temperature. The reetangular copper box B, which constitutes the warm temperature boundary, is cooled by liquid nitrogen from this outermost vessel E and is insulated with multilayer insulation from the outside. The whole apparatus is placed inside a vacuum vessel. A vacuum of at least 2 x 10-7 torr is maintained when measurements are made. The surface area of the 4.2-K fin is approximately 2.26 m2 (24.375 te) . The total heat absorbed by the 4.2-K fin is divided by this area to give the experimental heat transferrate per unit area between the 77- and 4.2-K surfaces. Large storage volumes are provided for both liquid helium and liquid nitrogen so as to facilitate prolonged testing (15 to 20 days), which is essential for degradation and contamination studies. The pressure in the liquid nitrogen vessel can be reduced to produce a lower warm boundary temperature. Fig. 4. Exterior view of the thermal radiation test apparatus (without the insulation shroud and one side wall of the nitrogen box) with the large-diameter vacuum chamber in the background. 494 E. M. W. Leung, R. W. Fast, H. L. Hart, and J. R. Heim EXPERIMENTAL TECHNIQUE Instrumentation The temperature of B is monitored using several copper-Constantan thermocouples and that of A with calibrated carbon resistors. Variation of the temperature across both A and B is calculated and measured to be negligible. The boil-off rate of the liquid helium in Cis measured with a calibrated wet test meter (accuracy ±0.2% ). An electrical signal is sent to a running strip chart recorder upon completion of each revolution of the meter indicator suchthat continuous (and therefore average) readings can be taken. The temperature of the evolving gas is read directly on the wet-meter and the liquid helium boil-off rate at 273.16 K and 760 mm Hg can then be calculated. It is important to note that the ratio of the saturated liquid density to the helium vapor density at 4.2 K is only 7.5:1 at 1 atm; therefore, the actual heliumevaporationrate should be 1.154 times the boil-off rate measured with the wet-meter. A typical experimental run produces results similar tothat shown in Fig. 5. The loss rate measurement is influenced by the variation of atmospheric pressure, which is therefore monitored continuously. A correction C4 ] for the rate of vaporization resulting from small atmospheric pressure changes can be approximated by the general formula dmr = !Vdp (4) dt dt and in the case of liquid helium, ER= 7.9 x 10- 3 V (dt) (4') 40.0 30.0 e ~ 20.0 E .5 ~ ·0 10.0 Ist 2nd 3rd 4th 5th 6th 7th 81h 9 th IO ih lllh nme . Days Fig. 5. Typical data obtained from a run. Run No. 10: Eight layers of NRC-2 on blackened 4.2-K surface, Eccoshield Al tape on 77-K surface. Average heat Ioad into the 4.2-K surface per unit area Q/ A = 13.45 ± 0.5 mW /m 2 • Teebniques for Redudug Radiatiou Heat Trausfer betweeu 77 aud 4.2 K 495 This explains the ±5 to 10% variation in the boil-off rate data. Forthis reason, each run is generally conducted over a period of a week or more to average out the effect of pressure variations and to insure the achievement of equilibrium conditions. Callbration The only major heat source into Cis heat conduction down the vent tube F. This was calculated tobe 8.3 mW using the method described in Scott [16]. Background runs of the apparatus without both A and B were performed. In the beginning, thermal acoustic oscillation was encountered and a background heat leak value five times higher than calculated was observed. The thermal acoustic oscillation was removed by filling the vent tube with cotton plugs which separated the gas ftow into small ftow channels C7 ]. The background run then yielded a result of 10.7 mW. The Iiterature on thermal acoustic oscillations C8 - 20 ] revealed that the system was nearly stable against oscillation and at the expected boil-off rates the background heat leak down the vent tube would only be about 0.3 mW. A sample run with the Eccoshield aluminum tape on both the 4.2- and 77-K surface gives a heat transferrate per unit area of 1.40 ± 0.05 mW /fe, which compares weil with the value obtained by Chaussy [9 ]. Six months later, after six other experimental runs, this configuration was tested again. A result of 1.38 ± 0.05 mW /fe was obtained, showing good reproducibility. lt is estimated that the overall accuracy of the measurements are ±10%. RESULTS AND DISCUSSION Aseries of six runs were performed using NRC-2 on a 4.2-K surface painted with 3M black Velvet paint [1 5 ] to simulate a dirty stainless steel helium vessel. Eccoshield tapewas applied to the 77-K surface. The results are shown in Fig. 6. Other data are given in Table I. By comparing the results of runs Nos. 1, 3, and 4, it can be seen that if in the NRC-2 series, one had used bare ETP copper at 77 K, the heat ftux would have been higher and vice versa for the 3M Al tape. 18 17 17.1 mW/m 2 16 NE ~ E .!: 15 ~ ·0 14 13 12 L---~~o----'zo----~-----~r---~5o.---·ro--~ro~--~aro No. of Layers of NRC -2 Fig. 6. Radiation heat ftux vs. number of layers of NRC-2 in a 50.8-mm gap from 77 to 4.2 K. E. M. W. Le-., R. W. Fut. H. L.a.rt. ud J. R. Heilll 496 Table I. Heat Transfer from 77.4 to 4.2 K System Run No. 4.2-K surface 77-K surface 1. Eccoshield Al tape Eccoshield Al tape 2. Eccoshield Al tape Eccoshield Al tape 3. Electrolytic Tough Pitch Cu (rinsed with Bright Dip and cleaned with Oakite) Tough Electrolytic Pitch Cu (rinsed with and Dip Bright cleaned with Oakite) 4. 3M Al tape 3M Al tape 5. Blackened with black 3M Velvet paint; then 8 Jayers of NRC-2 wrapped around the black surface 6. Blackened with black 3M Velvet paint; then 8 Jayers of NRC-2 wrapped around the black surface Blackened with black 3M Velvet paint; then 8 Jayers of NRC-2 wrapped around the black surface 1 layer of double-side Mylar aluminized epoxied to 77-K CU surface using Crest # 7344 epoxy* 6 layers of NRC-2 attached by nylon screws and polyester threads. ftoating Aluminum shield 7. Heat transfer rate unit area, mW/m2 (mW/ft2 ) Remarks Calculated value is 16.5 mW/m2 using emissivities values deduced from data of Chaussy et al. 14.9 ± 0.5 Same surfaces after being (1.38 ± 0.05) left in Iab and in apparatus for 6 months. No degradation observed. Repeatibility of data checked. 15.6 ± 0.8 Experiment was run for 11 (1.45 ± 0.075) days continuously. No sign of degradation observed. Typical method used on Jarge magnets previously constructed at Fermilab. # 425 lndustrial Tape 12.4 ± 0.5 (1.15 ± 0.05) seems to have better adhesive property at 77-K than Eccoshield Tape. Aluminized Mylar, mea28.8 ± 0.5 sured to have a 500 A alu(2.00 ± 0.05) minum Jayer on each side. t Outgassing of epoxy detected. 17.3±0.5 77-K surface has 2 inter(1.60 ± 0.05) weaving areas carefully prepared. 15.1 ± 0.5 (1.40 ± 0.05) 18.1 ± 0.5 Nylon screws with stainless (<1.68 ± 0.05) steel nuts were used as the standoffs. Floating shield made from 0.127-mm-thick aluminum sheets. * Crest Products Company, Santa Ana, California. t King-Seeley Company, same supplier as NRC-2. The measured value from run No. 1 (15.1 mW/m 2 ) is well approximated by (1), assuming e 77 = 0.03 and E4.2 = 0.011 (16.5 mW/m 2 ). However, the geometry of the apparatus, a ftat plate inside a reetangular box, may not be adequately represented by (1). A detailed anal~is using configuration factors gives a heat transferrate per unit area of 17.2mW/m2 , using the emissivities above. This indicates that geometrical effects account for 4 to 5% of the difference. Residual resistivity ratio (RRR) measurements of a few samples were made and are included in Table II. This gives a relative indication of how good the sample is as an infrared re