Loss-of-Flow Analysis of an Unfinned, Graded Fuel
Meat, LEU Monolithic U-10Mo Fuel Design in
Support of the MITR-II Fuel Conversion
by
Sarah M. Don
Submitted to the Department of Nuclear Science and Engineering
in partial fulfillment of the requirements for the degrees of
Master of Science in Nuclear Science and Engineering
MASSACHUSETTS IN6TlJTE,
and
OF TECHNOLOGY
Bachelor of Science in Nuclear Science and Engineering
at the
MASSACHUSETTS INSTITUTE OF TECHNOLOGY
OCT 2 9 201
LIBRARIES
September 2014
Massachusetts Institute of Technology 2014. All rights reserved.
Signature redacted
-..................
Department of Nuclear Science and Engineering
Author ...............
August 8, 2014
Signature redacted
.........--------Lin-wen Hu, Ph.D.
C ertified by ...............
Associate Director of Research Development and Utilization,
Nuclear Reactor Laboratory
Thesis Supervisor
Signature redacted
Certified by............
Benoit Forget, Ph.D.
Associat6#rofessor of Nuclear Science and Engineering
Thesis Reader
.
Signature redacted
.
Accepted by....
uji
azimi, Ph.D.
TEPCO Professor of uclear Engineering
Chair, Department Committee on Graduate Students
2
Loss-of-Flow Analysis of an Unfinned, Graded Fuel Meat,
LEU Monolithic U-10Mo Fuel Design in Support of the
MITR-II Fuel Conversion
by
Sarah M. Don
Submitted to the Department of Nuclear Science and Engineering
on August 8, 2014, in partial fulfillment of the
requirements for the degrees of
Master of Science in Nuclear Science and Engineering
and
Bachelor of Science in Nuclear Science and Engineering
Abstract
In order to satisfy requirements of the Global Threat Reduction Initiative (GTRI),
the 6 MW MIT Research Reactor (MITR-II) is to convert from the current 93%-enr
235
U highly-enriched uranium (HEU) fuel to the low-enriched uranium (LEU, <20%
235
U) fuel. This reduction in enrichment decreases the neutron flux level due to parasitic absorption by 238U. The neutron flux may be compensated for by increasing the
reactor's nominal operating power level to 7.0 MW. Thus a neutronic and thermalhydraulic study was undertaken to evaluate new fuel designs with graded fuel meat
thickness and unfinned clad that provide sufficient safety margins for steady-state
operation at 7.0 MW.
A previously-studied 18-plate LEU fuel design and an identical unfinned fuel design were compared to evaluate the effect of fin removal, demonstrating the need for
fuel redesign. A recent feasibility study has shown that a 19-plate, unfinned fuel design
with graded fuel meat thicknesses (19B25) provides fuel cycle length and steady-state
thermal hydraulic safety margins that meet the design criteria. The objective of this
study was to use the RELAP5 MOD3.3 code to confirm the steady-state thermalhydraulic safety margin and to analyze the loss-of-flow (LOF) transient performance
of this candidate fuel design.
Power distributions obtained for beginning-of-life (BOL), middle-of-life (MOL),
and end-of-life (EOL) were analyzed to study the effect of core power distribution
and burnup-dependent thermal properties on safety margins. Results show that the
MITR-II can safely operate at 7.0 MW with the proposed LEU fuel with an adequate
margin (40%) to the onset of nucleate boiling (ONB) -limiting power level.
The
minimum margin between coolant channel wall and saturation temperatures was at
3
least 22 C in the most limiting channel, in the most limiting core (BOL) at 7.0 MW.
The proposed LEU fuel design also performed well during a simulated LOF transient
after operation at 7.0 MW, with a peak fuel temperature of 106 C reached in the
hot channel, which is well below the U-1OMo blistering temperature of 365*C. During
the LOF transient, the maximum clad temperature was 980, meaning that no boiling
occurred even during the LOF transient. Bounding analysis to evaluate the effect
of an oxide layer and fuel meat thermal conductivity due to fuel burnup estimated
that up to a 15 C peak fuel temperature rise can be attributed to increased thermal
resistance of oxide layer and fuel thermal conduction reduction. Thus under the most
conservative assumption, the estimated peak fuel temperature is 121 C, well under
the blistering temperature limit of 365 C. It is concluded that the 19-plate unfinned
fuel design with graded fuel meat thickness is a promising candidate for the conversion
to LEU fuel and power uprate.
Thesis Supervisor: Lin-wen Hu, Ph.D.
Title: Associate Director of Research Development and Utilization,
Nuclear Reactor Laboratory
4
Acknowledgments
I would like to express my sincere appreciation of Dr. Lin-wen Hu and Dr. Tom Newton for their invaluable support and guidance throughout this project, and Professor
Benoit Forget for being my thesis reader.
Thanks also to my senior and mentor Eric Forrest who took me under his wing
providing me with the help and encouragement I needed to get started and stay on
track. Thanks to Dr. Koroush Shirvan for answering my RELAP questions late at
night and on weekends. Thanks to Dr. Erik Wilson and Dr. Floyd Dunn for providing me with the initial RELAP input decks and corresponding with me throughout
the project.
Thanks to Dr.
Kaichao Sun for supporting the neutronic aspects of
the project and providing me with neutronic data. Thanks also to Taylor Tracy for
assisting with the proof-reading of this paper.
I gratefully acknowledge the ANL/RERTR program for supporting this project.
This work was funded in part by the U.S. Department of Energy, Basic Energy Sciences, Office of Science, under contract BOA 2J-30101. This study is also sponsored
by the U.S. Department of Energy, National Nuclear Security Administration Office
of Global Threat Reduction.
Thank you to my family for being understanding that I cannot easily call or visit
often while we are on opposite sides of the world. Finally, I am eternally grateful to
my husband Maxwell Mann for supporting me through stressful times, even when he
is also incredibly busy as a graduate student.
5
6
Contents
1
2
Introduction
17
1.1
Global Threat Reduction Initiative . . . . . . . . . . . . . . . . . . .
18
1.2
Motivation for Conversion
19
1.3
High-Density Monolithic LEU U-10Mo Fuel
. . . . . . . . . . . . . .
21
1.4
Fuel Conversion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
22
1.5
Criteria for LEU Fuel Selection
23
. . . . . . . . . . . . . . . . . . . . . . . .
. . . . . . . .
. . . . . . . . . . .
Background
25
2.1
The MITR-II . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
25
2.1.1
Fuel .......
28
2.1.2
Absorbers . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
28
2.1.3
Experiment Facilities . . . . . . . . . . . . . . . . . . . . . . .
29
2.1.4
Cooling Systems
. . . . . . . . . . . . . . . . . . . . . . . . .
30
MIT Operating Limits . . . . . . . . . . . . . . . . . . . . . . . . . .
32
2.2.1
Onset of Nucleate Boiling
. . . . . . . . . . . . . . . . . . . .
33
2.2.2
Hot Channel Analysis
. . . . . . . . . . . . . . . . . . . . . .
34
2.2.3
Loss-of-Flow Transient Scenario . . . . . . . . . . . . . . . . .
36
2.2
2.3
MCODE ........
2.4
RELAP5 MOD3.3
................................
..................................
37
. . . . . . . . . . . . . . . . . . . . . . . . . . . .
37
2.4.1
Time Step Configuration . . . . . . . . . . . . . . . . . . . . .
38
2.4.2
Limitations . . . . . . . . . . . . . . . . . . . . . . . . . . . .
39
2.5
Previous LEU Fuel Design . . . . . . . . . . . . . . . . . . . . . . . .
43
2.6
High-Density Monolithic LEU U-10Mo Fuel
44
7
. . . . . . . . . . . . . .
3
2.6.1
Material Properties . . . . . . . . . . . . . . . . . . . . . . . .
44
2.6.2
Fuel Temperature Limit . . . . . . . . . . . . . . . . . . . . .
44
2.6.3
Graded Monolithic LEU U-10Mo Fuel Meat
47
. . . . . . . . . .
Research Objectives
51
4 Effect of Fin Removal
4.1
4.2
5
53
MIT27 Reference Case . . . . . . . . . . . . . . . . . . . . . . . . . .
54
4.1.1
Steady-State Analysis
. . . . . . . . . . . . . . . . . . . . . .
57
4.1.2
Loss-of-Flow . . . . . . . . . . . . . . . . . . . . . . . . . . . .
59
MIT30 - The Effect of Removing Fins . . . . . . . . . . . . . . . . . .
59
Proposed New Fuel Design
65
5.1
Fuel Cycle and Core Power Distributions . . . . . . . . . . . . . . . .
66
5.2
Axial Power Profiles
70
5.2.1
. . . . . . . . . . . . . . . . . . . . . . . ....
Simulated Conditions . . . . . . . . . . . . . . . . . . . . . . .
77
5.3
Beginning-of-Life
. . . . . . . . . . . . . . . . . . . . . . . . . . . . .
78
5.4
Middle-of-Life . . . . . . . . . . . . . . . . . . . . . . . . . . . ... . .
83
5.5
End-of-Life
86
5.6
ONB-Limiting Power Level . . . . . . . . . . . . . . .
. . . . . . .
89
5.7
Summary of RELAP Results . . . . . . . . . . . . . . . . . . . . . . .
90
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
6
2D Semi-Analytical Validation
7
Burnup Effect on Fuel Temperature
101
7.1
Effect of Oxide Layer . . . . . . . . . . . . . . . . . . . . . . . . . . .
102
7.2
Effect of Fuel Thermal Conductivity
105
8
93
Summary and Conclusions
. . . . . . . . . . . . . . . . . .
109
8
List of Figures
MITR-II Cutaway . . . . . . . . . . . . . . . . . . . . . . . . . . . .
26
2-2
MITR-II Core Layout . . . . . . . . . . . . . . . . . . . . . . . . . .
27
2-3
Photograph of MITR-II fuel element
. . . . . . . . . . . . . . . . .
27
2-4
MITR-II Cooling Systems . . . . . . . . . . . . . . . . . . . . . . .
31
2-5
MITR-II Natural Circulation . . . . . . . . . . . . . . . . . . . . . .
31
2-6
Typical forced convection subcooled boiling curve. [12]
. . . . . . .
34
2-7
Pump Coastdown Curve . . . . . . . . . . . . . . . . . . . . . . . .
36
2-8
RELAP Laminar Flow Treatment . . . . . . . . . . . . . . . . . . .
41
2-9
Fuel plate layer configuration. (Not to scale) [6]
45
.
.
.
.
.
.
.
.
2-1
.
. . . . . . . . . . .
2-10 Optical microscopy images of the U-IOMo fuel-Zr diffusion barrier in-
48
4-1
MIT27 Nodalization Diagram
4-4
MIT27 Finned Fuel Steady-State Temperatures
. . . .
57
4-2
MIT27 Plate/Channel Stripe Configuration . . . . . . .
58
4-3
.
. . . . . . . .
45
MIT27 Power Profiles . . . . . . . . . . . . . . . . . . .
58
4-5
MIT27 Finned Fuel Transient Temperatures and Flow.
.
2-11 End plate heat flux reduction by fuel meat gradation
.
.
terface. [4] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
60
4-6
Effective Gap After Fin Removal
. . . . . . . . . . . .
61
4-7
MIT30 Steady-State Axial Temperature Profiles . . . .
62
4-8
ONB-Limiting Power Level for MIT27 and MIT30 Fuels
62
5-1
19B25 RELAP structure labeling . . . . . . . . . . . .
66
5-2
19B25 Nodalization Diagram . . . . . . . . . . . . . . .
67
.
.
.
.
.
.
.
. . . . . . . . . . . . . .
9
55
5-3
Power generation shift from BOL to EOL .6
. . . . . .
69
5-4
BOL Power Profiles . . . . . . . . . . . . . . . . . . . . . . . . . . . .
71
5-5
MOL Power Profiles
. . . . . . . . . . . . . . . . . .
73
5-6 EOL Power Profiles . . . . . . . . . . . . . . . . . . . . . . . . . . . .
75
5-7
Fractional power and flow after the LOF scram
. . . . . . . . . . . .
79
5-8
BOL Steady-State Temperature Profiles
. . . . . . . . . . . . . . . .
81
5-9
19B25 BOL LOF Transient Temperatures
. . . . . . . . . . . . . . .
82
. . . . . . . ..
5-10 MOL Steady-State Temperature Profiles . . . . . . . . . . . . . . .
84
5-11 19B25 MOL LOF Transient Temperatures
. . . . . . . . . . . . . . .
85
. . . . . . . . . . . . . . . .
87
. . . . . . . . . . . . . . .
88
. . . . . . . . . . . . . . . . . . .
90
6-1
Semi-analytical model temperature profiles . . . . . . . . . . . . . . .
98
7-1
Fuel Temperatures With Oxide
. . . . . . . . . . . . . . . . . . . . .
104
7-2
Thermal Conductivity of U-10Mo . . . . . . . . . . . . . . . . . . . .
106
5-12 EOL Steady-State Temperature Profiles
5-13 19B25 EOL LOF Transient Temperatures
5-14 19B25 ONB-Limiting Power Levels
10
List of Tables
1.1
Research reactors that require high-density fuel to convert to LEU. [30]
20
1.2
Summary of HEU and LEU fuels for the MITR-II. [13]
21
2.1
Summary of primary loop dimensions in RELAP input. [13]
2.2
Reference time step configuration [2]
. . . . . . . .
. . . . .
39
. . . . . . . . . . . . . . . . . .
40
2.3
Optimal time step configuration . . . . . . . . . . . . . . . . . . . . .
40
2.4
Composition of LEU U-10Mo monolithic fuel. [2]
. . . . . . . . . . .
45
2.5
Thermal properties for Al-6061. [10]
. . . . . . . . . . . . . . . . . .
46
2.6
Thermal properties for zirconium. [10]
. . . . . . . . . . . . . . . . .
46
2.7
Thermal properties for LEU U-Mo. [10] . . . . . . . . . . . . . . . . .
46
2.8
Graded fuel meat thickness combinations. [2] . . . . . . . . . . . . . .
49
4.1
MIT27 Fuel Geometry . . . . . . . . . . . . . . . . . . . . . . . . . .
56
4.2
MIT27 and MIT30 ONB-Limiting Power Levels . . . . . . . . . . . .
61
4.3
MIT30 Fuel Geometry . . . . . . . . . . . . . . . . . . . . . . . . . .
63
5.1
19B25 Fuel Geometry . . . . . . . . . . . . . . . . . . . . . . . . . . .
68
5.2
BOL Hot Channel Factor and Power Per Plate . . . . . . . . . . . . .
72
5.3
MOL Hot Channel Factor and Power Per Plate
. . . . . . . . . . . .
74
5.4
EOL Hot Channel Factor and Power Per Plate . . . . . . . . . . . . .
76
5.5
19B25 steady-state conditions . . . . . . . . . . . . . . . . . . . . . .
79
5.6
LOF Maximum Temperature Rise . . . . . . . . . . . . . . . . . . . .
83
5.7
19B25 Hot Channel Maximum Temperatures . . . . . . . . . . . . . .
91
11
6.1
Comparison of RELAP results with semi-analytical model for 19B25
B O L.. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
99
7.1
Oxide Sensitivity
7.2
Burnup Effect on Thermal Conductivity
. . . . . . . . . . . . . . . .
107
7.3
Burnup Effect on Peak Fuel Temperature . . . . . . . . . . . . . . . .
107
. . . . . . . . . . . . . . . . . . . . . . . . . . . . .
12
103
Nomenclature
Abbreviations
19B25
19-plate unfinned graded LEU fuel meat fuel design
ACI
Advanced clad irradiation facility
Al-6061
6061 alloy of aluminum
ANL
Argonne National Laboratory
ASV
Anti-siphon valve
BNCT
Boron neutron capture therapy
BOC
Beginning of (fuel) cycle
BOL
Beginning of (core) life
CHF
Critical heat flux
enr
23
1U
enrichment
EOC
End of (fuel) cycle
EOL
End of (core) life
gpm
gallons per minute
GTRI
Global Threat Reduction Initiative
HEU
High enriched uranium
HTIF
High temperature irradiation facility
ICSA
In-core sample assembly
INL
Idaho National Laboratory
LEU
Low enriched uranium
LOF
Loss-of-flow
LSSS
Limiting Safety System Setting
LWR
Light water reactor
MCODE
MCNP-ORIGEN Depletion Program
MIT27
Reference 18-plate LEU finned fuel design
MIT30
18-plate LEU unfinned fuel design
13
MITR-II
Massachusetts Institute of Technology Research Reactor II
MOC
Middle of (fuel) cycle
MOL
Middle of (core) life
MWt
Megawatts thermal energy
NCV
Natural circulation valve
NRC
Nuclear Regulatory Commission
OFI
Onset of flow instability
ONB
Onset of nucleate boiling
OSV
Onset of significant voiding
RELAP
Reactor Excursion and Leak Analysis Program
RERTR
Reduced Enrichment for Research and Test Reactors
RIA
Reactivity Insertion Accident
SAR
Safety Analysis Report
U-10Mo
Uranium fuel with 10 molybdenum atoms for every uranium atom
UAlx
Uranium-aluminum alloy
Symbols
Af
Flow area
cp
Specific heat (kJ/kg K)
De
Hydraulic diameter
hc
Heat transfer coefficient (W/m2K)
k
Thermal conductivity (W/mK)
rh
Mass flow rate (kg/s)
p
Viscosity (Pa s)
nch
Number of channels
Nu
Nusselt Number
P
Pressure
Pw
Wetted perimeter
14
Prandtl Number
q"
Heat flux (W/m2
)
Pr
)
Volumetric heat generation rate (W/m 3
p
Density
Re
Reynolds Number
Tclad,ONB
ONB cladding temperature ( C)
Tsat
Saturation temperature ( C)
w
Width
Subscripts
ci
Clad inner surface
CL
Centerline (fuel)
co
Clad outer surface
f
Fuel
g
Channel gap
0
Oxide layer
p
Fuel plate
w
Water
z
Zirconium layer
15
16
Chapter 1
Introduction
The MITR-II is a 6.0 MW research reactor currently fueled with 93%-enr
Fuel element design and feasibility studies are ongoing for low-enriched
235
235
U fuel.
U (LEU)
fuel conversion in accordance with the Global Threat Reduction Initiative (GTRI).
LEU fuel is defined as containing uranium enriched to < 20%
235
U. A prime objec-
tive of the GTRI is to convert all test and research reactors worldwide to LEU fuel.
Research reactors typically use highly-enriched
2 35
U (HEU) fuel to achieve a high
neutron flux in a compact core for applications such as neutron beam experiments,
isotope production, and accelerated materials irradiation testing. The objective of
removing HEU fuel from reactors is to reduce the global weapons proliferation threat
by reducing the inventory of weapons-grade nuclear materials. The MITR-II is one of
the six research and test reactors in the U.S. that require a new, high-density fuel in
order to convert to LEU. A decrease in enrichment reduces the neutron flux density
due to parasitic absorption in
23
8U. [19] However, this reduction in reactor perfor-
mance may be countered by a power uprate to 7.0 MW.
Several new fuel geometries have been proposed in recent studies for the MITR-II
conversion. [19, 9, 13, 14] The Reduced Enrichment for Research and Test Reactors
(RERTR) program has been developing high-density alternate fuel meats to facilitate
the conversion. U 3Si 2 and U-1OMo dispersion-type fuels have been developed for LEU
reactor fuel, though some reactors, including the MITR-II, can only convert to LEU
17
if the density is significantly higher than achievable in dispersion fuels. [22] Power
peaking in the current fuel geometry would prevent a power uprate of the MITRII, and so new geometries with uniform radial power distributions have also been
designed and assessed by the GTRI program and MIT. The most promising design
to date is a 19-plate, unfinned, high-density monolithic U-IOMo fuel with graded fuel
meat thickness (thinner in the outer six plates) that is referred to in this report as
the "19B25" case. The reference case to which this design is compared is an 18-plate,
finned, high-density monolithic U-10Mo fuel with the same fuel meat thickness in
every plate, which is referred to as "MIT27". [2] There is also an intermediate case
which is identical to MIT27 but with unfinned fuel plates called "MIT30" to study
the effect of fin-removal on the thermal-hydraulics of the MITR-II core.
Research
concerning the neutronic aspects of new fuel designs is currently underway in parallel
with thermal-hydraulic studies to support the LEU fuel design, core conversion and
power uprate.
1.1
Global Threat Reduction Initiative
The mission of the GTRI program is to reduce the amount of vulnerable nuclear
materials, which involves the conversion of HEU-fueled reactors to LEU, removal of
excess nuclear materials, and protection of nuclear materials from theft. The GTRI
program goals announced by Energy Secretary Abraham on May 26, 2004, included
the conversion of most civilian reactors that use HEU to LEU by 2014. In order to
allow time for reactor conversion, fuels suitable for conversion had to be developed
well in advance of this date. Considering this, supplying qualified fuels required for
conversion of most research reactors to LEU by the end of 2010 was initially planned.
[30] MITR-II conversion has been postponed due to the delay in high-density LEU
fuel qualification.
The MITR-II currently employs robust security measures and "just-in-time" receipt and storage of nuclear materials, though it still requires the manufacture, trans-
18
port, use and storage of HEU fuel. The core must be converted to LEU fuel in order to
comply with 10 CFR 50.64 requirements listed below so that it can continue operating
when a suitable LEU fuel becomes available.
1.2
Motivation for Conversion
RERTR is a global program whose objective is to support research and development
of high-density fuels that can be used to convert HEU-fueled research reactors to
LEU-fueled. It is an effort to promote defense and security by reducing the risk of
theft or diversion of HEU-fuel. The successful conversion of US non-power reactors to
LEU fuel should encourage similar action by reactor facilities globally, thereby reducing the amount of HEU fuel at international facilities. In 1984 the NRC proposed the
rule requiring all new licensees to use LEU fuel and all existing licensees to convert
from HEU to LEU, when suitable fuel becomes available. Funding for the conversion
efforts is provided by the federal government through the GTRI program.
The Ford Nuclear Reactor (FNR) at the University of Michigan was the first reactor converted by the RERTR program in December 1981 using a UAl'-Al LEU fuel
with a density of 1.7 g/cm 3 , but most other reactors require LEU fuels with significantly higher uranium density.
[30] The uranium density proposed for the MITR-II
conversion is 15.3 g/cm3 . [13] Table 1.1 lists the research reactors that need(ed) highdensity fuel in order to successfully convert to LEU fuel. There is no set deadline for
completion of conversion efforts, however there is limited supply of HEU for research
reactors remaining. [20]
The MITR-II and other HEU-fueled research reactors are currently supplied with
fuel containing recycled uranium from dismantled Russian nuclear warheads as part of
the Megatons to Megawatts program established in 1994. As of the end of 2013 when
the contract expired, approximately 20,000 Russian warheads had been dismantled
as part of this effort. [28]
19
Table 1.1: Research reactors that require high-density fuel to convert to LEU. [30]
Country
Reactor
Power (MW) HEU Consumption (kg/yr)
Belgium
BR2
80
29
RHF
57
55
ORPHEE
JHR
14
100
16
-
France
Germany
FRM-II
USA
MITR-II
MURR
NBSR
HIFR
ATR
ATRC
20
6
38
>5
10
24
20
100
13
80
250
0.005
120
0
10
2.5
14
8
2
6
15
6
21
18
15
15
9
100
62
0
0
Russia
LWR-15
IRT-MEPI
IR-8
IRT-T
VVR-TS
VVR-M
IVV-2M
MIR-Mi1
CAMIR-M 1
1
The MITR-II has two main requirements for the conversion. A high-density fuel
meat material must be available, and a new fuel geometry that allows operation up
to 7.0 MW must be designed with adequate steady-state safety margins to onset of
nucleate boiling (ONB). The RERTR program supports research and development of
high-density monolithic LEU U-1OMo fuel that is proposed for the MITR-II conversion. RERTR and MIT are collaborating to perform neutronic and thermal hydraulic
analyses of the proposed fuel designs.
20
1.3
High-Density Monolithic LEU U-10Mo Fuel
The proposed LEU fuel type is a monolithic (non-dispersion) U-10Mo material with
235
U enriched to 19.75%. Studies at MIT have shown that high-density U-10Mo fuel
with a uranium density of at least 15 g/cm 3 is the only feasible LEU fuel option for
the MITR-II. [13, 29, 19] The facility is awaiting the successful qualification of this
high-density LEU fuel in order to proceed with the conversion. [13]
Preliminary RERTR experiments showed that the presence of an interaction layer
between the fuel and cladding materials caused mechanical internal stress problems.
To minimize the fuel-cladding interaction, introducing a diffusion barrier between the
cladding and fuel was proposed. The current monolithic plate design includes a 0.0254
mm thick, 99.8% pure, annealed zirconium diffusion barrier between the U-IOMo fuel
and the Al-6061 clad. Table 1.2 provides a summary of current HEU and proposed
LEU fuel properties. Miniature plates of this U-lOMo-Zr-Al-6061 configuration were
irradiated in the Advanced Test Reactor (ATR) at Idaho National Laboratory (INL)
with promising irradiation performance. [21]
The HEU fuel failure limit is the Al-6061 softening temperature of 450 C. For
the LEU fuel, however, the limiting factor was found to be blistering caused by
temperatures exceeding 365 C. [4] Verification that the fuel temperature does not
exceed 365 C during a loss-of-flow accident is necessary in order to qualify this LEU
fuel for the MITR conversion.
Table 1.2: Summary of HEU and LEU fuels for the MITR-II. [13]
Parameter
HEU
LEU
Fuel meat composition
U-Alt U-10Mo
2 35
U enrichment (%)
93
19.75
Uranium density (g/cm 3 )
Mass of 235 U per element (g)
Heat capacity (at 100 C, J/kg C)
Melting point ( C)
21
1.54
15.3
508
627
968
142
1400
1135
1.4
Fuel Conversion
A report produced by ANL in July 2013 presents a candidate fuel design that uses
18 finned fuel plates, thinner cladding, thinner coolant channels, and high density
uranium-molybdenum monolithic alloy fuel enriched up to 19.75%
2 35U.
Dunn et al.
[9] Transient analysis studies for reactivity insertion accident (RIA) and loss of flow
(LOF) scenarios showed that the maximum fuel temperature would not reach the
fuel blistering temperature of approximately 365 C, which makes these fuel designs
candidates for the fuel upgrade. [9]
A previous study showed that a design containing 18 fuel plates with 0.508 mm
thick U-10Mo LEU fuel with 0.25 mm finned cladding enables the MIT reactor to
retain its current flux with a power uprate to 7 MW while leaving sufficient margins
to ONB. [18] More recent studies at ANL suggest several designs with 18-19 unfinned
fuel plates, various fuel-to-clad thickness ratios, minimum required power uprate to
6.7 MW to 7.1 MW to maintain flux, and ONB-limiting power levels of 8.6 MW to
10 MW. [9, 2]
The latest candidate fuel designs have graded fuel meat. Currently the hot channel
(most limiting) in the MITR-II core is the end channel (for even fuel meat thickness
across plates) due to extra moderation. In order to reduce this peaking in the end
channel, the design suggests that outer 2-3 fuel plates have graded, thinner fuel meats
than the inner plates, creating a more even temperature profile across the fuel element, reducing the peak-to-average fuel temperature ratio and increasing the limiting
power. This also makes it more difficult to identify the hot channel, so several channels must be studied to ensure the hot channel is studied. [9]
22
1.5
Criteria for LEU Fuel Selection
In order for a fuel design to be feasible for the LEU conversion, it must meet the established safety criteria for the Safety Analysis Report (SAR). The new fuel must have
negative void and temperature coefficients, adequate shutdown reactivity margins at
all stages of core life, sufficient excess reactivity to overcome parasitic absorption during
135
Xe transients, be adequately cooled so as to avoid ONB in the hot channel,
and allow natural circulation to be effective at decay heat removal upon shutdown.
In order for the fuel to improve operating characteristics it must generate a thermal
and/or fast flux equal to or greater than that of the HEU core at the same power
level, and the fuel cycle length must be equal to or longer than that of the current
HEU core at the same power level.
23
24
Chapter 2
Background
This study is motivated by the need for a new fuel for the MIT research reactor, and
a power uprate to maintain the current neutron flux essential for the experimental
facilities. This section gives a brief overview of the reactor and experimental facilities,
thermal-hydraulics, the codes used, and previous work towards the MITR LEU fuel
conversion.
2.1
The MITR-II
The MITR-II is a research reactor licensed by the NRC for 6.0 MW operation until
2030.
The reactor is nominally operated at 5.9 MWt, well within the window of
permissible operating conditions as determined by the Safety Analysis Report (SAR).
[24] Figure 2-1 is a cutaway of the MITR-II showing the core situated in the core tank
and the various experiment access ports. The light water cooled and moderated tanktype reactor is heavy water and graphite reflected. Figure 2-2 shows the reactor core
which contains 27 fuel element positions, interior fixed absorbing plates, 6 control
blades and one fine control regulating rod. Typically only 24 positions contain fuel
elements while 3 positions are occupied by experiments or dummy elements.
The
reactor provides a high thermal neutron flux environment for the generation of medical
isotopes and in-core experimental studies. It also facilitates education through the
student operator training program, tours and reactor safety and technology courses.
25
5
Core Tank
Concrete
Pneumatic
Experiment Port
-
Shielding
rRe
Core
D20
Reflector
-
Figure 2-1: MITR-II Cutaway
26
Medical
Treatment Room
Fuel Element
I
C-12
Gc11
\
/
B-6
A-2
C-7
Core Tank
C-4
\
B-9
C-15
AB-1
C-1
B-4N
Absorbers
\4/Fixed
c-
C-2
B-g2
63
us
\
Regulating Rod
A-1
/
N7
C-13\
B-S
B-7
A-3
\t
C-8
'
c4
ad
C3
Control made
2
Figure 2-2: MITR-II Core Layout
Figure 2-3: Photograph of MITR-II fuel element
27
2.1.1
Fuel
Each fuel element is comprised of 15 aluminum-clad fuel plates milled with fins 0.01
inches wide and 0.01 inches tall as shown in Figure 2-3. All fuel plates are identical
with fuel meat measuring 0.2 inches thick with a frame of 0.03 inches of aluminum
cladding to the sealed edges. The fuel meat is a U-Al, cermet which can withstand
high temperatures like a ceramic while displaying ductility like a metal. The aluminum cladding softens at approximately 450 C which is considered to be the failure
limit for the fuel. [24] The rhomboid shape of each fuel element facilitates more even
and efficient fuel burnup by permitting 1800 rotation and inversion.
Currently the MITR achieves fast and thermal fluxes of up to 1.2 x 1014 n/cm2 s
and 3.6 x 1013 n/cm2 s respectively, with 93%-enr fuel at 5.9 MWt. It is estimated that
with conversion to 19.75%-enr high-density monolithic U-10Mo fuel that the neutron
flux at 5.9 MWt will be reduced by 10-20% due to parasitic neutron absorption in
23
8U.
[19]
2.1.2
Absorbers
The MITR-II core contains 6 control blades, a regulating rod, and fixed absorbers.
The absorbers permit precise reactivity control and power shaping.
The control blades hug each side of the hexagonal core, as shown in Figure 2-2.
Each blade is attached to a drive mechanism by an electromagnet, and is made up
of two 1.1% boron-impregnated stainless steel blades 0.125" (3.175 mm) thick with a
0.05" (1.27 mm) gap to allow for cooling. Only one shim blade can be moved outward
at a time to limit the rate of reactivity insertion. During startup and steady-state
operation, the six shim blades at kept at the same position as a bank to prevent
radial power peaking. When a scram condition occurs, current to the electromagnets
holding the blade to their drive mechanisms is cut, and the control blades drop back
into the core, shutting down the reactor in less than one second. [15]
28
The cadmium-lined fine control regulating rod is located at one of the six corners
outside the core housing, and is connected to the auto-control system to keep power
steady. When the reactor is on auto-control, the regulating rod is automatically inserted and withdrawn as needed to compensate for small reactivity changes due to
moderator temperature, coolant temperature, and xenon effects. When the regulating rod works its way outside of its useful range (too far in or out) the operator drives
the regulating rod back to a more reactive position and compensates for the reactivity
change with the control blade bank.
The fixed absorber is made up of several plates that are fixed in the core lattice
as shown in Figure 2-2. The 1.1% boron-impregnated stainless steel plates reduce the
radial power peaking in the center of the core.
2.1.3
Experiment Facilities
The MITR-II has experiment facilities both inside and outside the core for medical,
activation and materials performance studies.
In-core experiments are loaded into
positions Al, A3 and B3 (see Figure 2-2) for fast-flux irradiation. An in-core sample
assembly (ICSA), which is typically loaded into position A3, is capable of reaching
temperatures up to 850 C. [16, 2] The Advanced Clad Irradiation (ACI) loop is heated
and pressurized to simulate typical power LWR conditions for accelerated clad ma-
terial irradiations. [16] The High Temperature Irradiation Facility (HTIF) achieves
temperatures of up to 1600 C with a fast flux of approximately 1 x 1014 n/cm2 s to aid
in the materials testing for next generation high temperature reactors. [16] Though
the ICSA, ACI and HTIF are used most frequently, the reactor is capable of irradiating various other kinds of assemblies for corrosion and fissile material studies.
Experiment facilities exterior to the core include several beam ports, vertical thimbles, pneumatic tubes, a gamma irradiation facility and a medical treatment room.
The beam ports service a diffractometer and a spectrometer for neutron scattering
29
and attenuation studies. The vertical thimbles and pneumatic tubes are used to place
a sample for irradiation as close to the core as possible for the highest neutron flux
external to the core. The gamma irradiation facility is provided by space in the spent
fuel pool where samples can be exposed to gamma radiation for extended periods
of time. The medical (fission converter) treatment facility was defueled in 2013 af-
ter cessation of the boron-neutron capture therapy (BNCT) trials for treatment of
glioblastoma multiforme tumors (for which the facility was built). With re-installation
of fuel and coolant the thermal/epithermal neutron beam could be utilized again for
experiments. [16, 15]
2.1.4
Cooling Systems
Figure 2-4 shows a simplified schematic of the MITR-II primary and secondary cooling
systems. The water that leaves the core tank is pushed through the main primary-
secondary heat exchanger by two identical primary pumps before returning to the
core tank. Water enters the core tank via the core inlet pipe and flows down the annular mixing (downcomer) region to the core inlet plenum at the bottom of the core
tank. The nominal primary coolant flow rate is 2000 gpm (125 kg/s) with the reactor
scram set to occur if flow drops below 1900 gpm (119 kg/s). The water is forced up
through the core between the fuel plates and into the mixing region above the core
before exiting the core tank through the outlet pipe. The core tank inlet and outlet
temperatures are nominally maintained at approximately 43 C and 52 C respectively.
A core bypass flow factor of 0.0795 was measured for the MITR-II, indicating less
than 8% of flow cools structures other than fuel plates. [24] Not shown in the diagram
is an ion column primary coolant clean-up system, another much smaller pump that
provides long-term decay heat removal and does not affect flow through the core,
and a primary water storage tank. The heat exchanger is a titanium plate-type heat
exchanger with opposing flow directions for the primary and secondary coolants.
On the secondary side, the water circulates from the main primary-secondary heat
30
Cooling
Towers
Pumps
Reactor
Heat
Secondary
Exchanger
Pumps
Figure 2-4: Simplified schematic of the MITR-II primary and secondary cooling systems. (Not to scale) [7]
r
Core
Core
outlet
Outlet
i
ASS
ASV
1
1
Core
Core
~Inlet
Inlet
1
1
1
i
1
1
1
NCV
NCV
1
1
1
Core
1
t
t
Core
1
y
l
Figure 2-5: Natural circulation paths are facilitated by the natural circulation and
anti-siphon valves. (Not to scale) [6]
31
exchanger to the cooling towers and back by forced convection provided by another
set of two identical secondary system pumps. As water evaporates from the cooling towers, the secondary system water inventory is replenished by city water. The
cooling tower outlet temperature is nominally maintained between 20*C and 30*C
depending on the season. Not shown in Figure 2-4 on the secondary side are the various auxiliary cooling systems (biological shield, experiments, A/C and heavy water
reflector) and their heat exchangers.
Natural Circulation
The MITR-II is designed for passive decay heat removal via natural circulation in the
core tank. There are four natural circulation valves (NCVs) located at the bottom of
the core tank and two anti-siphon valves (ASVs) located a the height of the coolant
inlet pipe, as shown in Figure 2-5. When forced convection is provided by the two
main primary pumps, the water forces the balls in the NCVs and ASVs up, sealing the
opening on the top of each valve. When the pumps stop and the flow is reduced, the
balls drop, allowing water to pass through the opening at the top, thus establishing
natural circulation.
Ten feet of water above the top of the fuel in the core tank
facilitates natural convection cooling. In this configuration, the water flows upwards
through the core as it is heated, and downwards around the periphery of the core as
it cools, permitted by the open NCVs and ASVs which maintain a natural circulation
loop.
2.2
MIT Operating Limits
Currently the Limiting Safety System Settings (LSSS) are based on ONB. The margin
between the LSSS and licensed steady-state power levels is 20%; the current LSSS
power level is 6.0 MW and the licensed power level is 7.2 MW. The licensed power
has an additional 20% margin to ONB.
32
2.2.1
Onset of Nucleate Boiling
A power uprate requires analytical confirmation that the reactor can operate within
the limit of nucleate boiling onset. Currently, the operating limits of the MITR-II are
based on the conservative Bergles-Rohsenow nucleate boiling correlation as shown
in Equation 2.1, where Tdad,ONB is the fuel cladding temperature (*C), Tat is the
saturation temperature ( C), q" is the local heat flux (W/m 2 ),
and P is pressure
(bar). [3, 24, 25] This nucleate boiling correlation was also used for the Japanese
research reactor JRR-3 which has thin rectangular vertical coolant channels similar to
the MITR-II. [24] Development of a correlation specific to the MITR-II fuel geometry
will aide in the development of fuel surfaces with a higher margin to the onset of
nucleate boiling (ONB) and thus provide for a power uprate.
ii
Tdad,ONB = Tsat
0.43Po.o234
+ 0.556 1082P1.156
2 1
The concern about ONB is that it is the initialization of nucleate boiling before
the point of critical heat flux (CHF). CHF is the condition at which the heat transfer deteriorates significantly, leading to elevated fuel temperatures.
Because CHF
is followed by a rapid rise in temperature, a phenomenon that occurs before CHF
can be used as the thermal-hydraulic limit so as to leave a margin before CHF. As
temperature increases in single phase flow, the first boiling phenomenon that occurs
is ONB (see Figure 2-6). This is when the cladding surface is able to cause bubble
formation, but the coolant is still subcooled so the bubbles do not detach from the
cladding surface. Following ONB is the onset of significant voiding (OSV) which is
the condition when bubbles are formed on the cladding surface, detach and travel
with the coolant.
OSV is closely followed by the onset of flow instability (OFI)
which is the condition when the flow rate decreases with void fraction increase. This
can lead to flow reduction and rapid temperature rise, quickly leading to CHF. [24, 31]
When the coolant channel flow rate is reduced, pressure drop decreases and eventually vapor is generated and void fraction increases. As flow is further reduced, there is
33
El
significant vapor build-up and the pressure drop increases with decreasing flow. When
the pressure and flow oscillate in this way it is characterized as OFI. This can lead to
significant flow reduction and rapid temperature rise, quickly leading to CHF. [24, 31]
CHF
Fully
Boiling
Partial
Boiling-,
ONB
m
OF
OSV
Single Phase
Heat Transfer
Tsar
Torre Tosv
Surface Temperatur, T.a,
Figure 2-6: Typical forced convection subcooled boiling curve.
[12]
Sporadic flow instability has been observed between primary pump failure and the
establishment of steady natural convection, though fuel temperature excursion
due
to flow instability has not been observed in any previous RELAP modeling
of the
MITR-II core. [13]
2.2.2
Hot Channel Analysis
The hot channel is a theoretical channel into which all the most limiting conditions
are combined; thickest oxide layer, lowest thermal conductivity,
lowest flow, highest power etc.
By analyzing the thermally limiting hot channel, it can be assumed
that the results obtained are the most conservative and envelope the worst
possible operating conditions and the highest temperatures.
This method is used in the
thermal-hydraulic analysis of a core before the failure, safety and operating
limits
34
are established. Steady state analysis of an unfinned, new type of LEU fuel must be
performed to support the fuel conversion study.
For the reference case (MIT27, see Section 4), the following conservative conditions
were lumped into one channel to engineer the hot stripe1 :
" The wetted perimeter is set equal to the heated perimeter so that the flow area
is reduced
" An oxide layer of 0.001 thickness displaces coolant flow area
" The heat transfer coefficient is reduced because of the oxide layer
* A flow disparity factor of 0.93 is taken into account by an adjustment in the hot
channel flow area. Flow disparity in this context means a difference (reduction)
of flow compared with other channels.
" The power profile is higher than average by a factor of 1.13
Not all these conditions were used in the 19B25 (see Section 5) hot channels because they were deemed to be excessively conservative. For example, no oxide layer
was included even on hot plates because it is a 22-element beginning of life core with
all fresh fuel that will only be run for 2-3 months before a mixed-burnup refueling is
used, and therefore the buildup of oxide during this time is negligible. For the 19B25
case, the following conservative conditions were lumped into one channel to engineer
the hot channel:
" The wetted perimeter is set equal to the heated perimeter so that the flow area
is reduced
" Temperature increase caused by surface oxide formation is negligible
'A hot "stripe" is a fraction of a wall or channel (an 1/8 th of a channel in this case). See Section
4 for a description of stripe handling for the MIT27 case.
35
* The power profile is higher than average by a factor of approximately 1.3 (it
varies with each type of hot channel, see Section 5.2).
2.2.3
Loss-of-Flow Transient Scenario
The accident scenario analyzed in this study was a LOF accident due to simultaneous failure of both primary pumps. This means that the reactor scrams at the
flow
scram point of 2200 gpm 2 which is more conservative than if the reactor scrammed
at a higher flow rate.
After the pumps fail, the system flow coasts down to zero
in approximately 10 seconds according to the pump coastdown curve in Figure
2-7.
After 5.4 seconds the anti-siphon and natural circulation valves automatically
open,
facilitating natural circulation in the core tank.
140,
120 2
*
"
HEU (1900 gpm)
100
*
0
0
0*
0
LEU (2200 gpm)
0
80
I
c"
60
oe
0
40
0
20
4
200
0
0
0.
0
2
2
4
*
0"
6
8
10
12
14
Time (s)
Figure 2-7: Coastdown curve for MITR-II primary pumps. [17]
Pump Coastdown
Measurements of the coastdown of the primary flow after loss of power to the primary
pumps were made on April 14, 2011 at the MITR-II. [17] That data was acquired for
2
The current MITR-II flow scram point but 1900 gpm and 2200 gpm is the proposed
scram point
for the new core since the new core will likely operate with a higher flow rate.
36
a starting flow of 2150 gpm. The proposed primary flow scram point for the 19B25
case is 2200 gpm. In order to use the pump coastdown curve in the RELAP model,
interpolation of the measured coastdown curve was used to develop a new curve for
2200 gpm. Figure 2-7 shows the new pump coastdown curve that was used in the
19B25 RELAP model (where 137 kg/s is equivalent to 2200 gpm at 50 C), compared
to the reference case pump coastdown (where 112 kg/s is equivalent to 1800 gpm at
50 C).
2.3
MCODE
MCODE (MCNP-ORIGEN Depletion Program) is a code developed at MIT that links
MCNP5 (Monte Carlo N-Particle code) with ORIGEN-2.2 (an isotope generation and
depletion code) in a user-friendly manner. MCNP is used to extract cross-sections
and flux values, and then the output is parsed into ORIGEN for depletion calculations. MCODE automates this interfacing, improving efficiency and reducing the
probability for introducing human error. MCODE was used by K. Sun in this study
to generate the power profiles for each type of fuel plate in three cores with varied
burnup, which were parsed into RELAP5 MOD3.3 for thermal-hydraulic analyses.
[26]
2.4
RELAP5 MOD3.3
RELAP (Reactor Excursion and leak Analysis Program) is a multidimensional thermalhydraulics and neutron kinetics modeling code for steady-state and accident transient
analysis of reactor cores developed by Idaho National Lab (INL) and distributed by
the Nuclear Regulatory Commission (NRC). [1] RELAP5 MOD3.3 is not the latest
version but it is the version that was available and for which MITR-II input decks had
already been benchmarked. [2] RELAP5 MOD3.3 has also been benchmarked against
37
RELAP5-3D up to the point of the onset of nucleate boiling (ONB) for several other
research reactors. [8] An MITR conversion-related study by Y. Ko found RELAP
to conservatively predict peak temperatures compared to MULCH (a multi-channel
thermal-hydraulics analysis code).
[14] The RELAP results obtained in this study
were also checked against a semi-analytical model.
This study employs the use of RELAP5 MOD3.3 to perform steady-state and
transient hot channel analyses for the MITR-II core with finned and unfinned fuel
plates. The input takes specification for flow circuit components (pumps, pipes
etc.),
initial conditions (flows, temperatures etc.), heat structures (fuel plates), boundary
conditions (which fuel plates deposit heat into which coolant channels), materials data
(heat transfer coefficients etc.), and power level. The individual axial power profiles
of each heat structure can also be specified. To run a transient case, additional information about the type of accident and what changes happen in the system during the
transient must be specified. For example, in the loss-of-flow (LOF) accident scenario
modeled in this study, the accident begins with loss of off-site power which trips the
main pumps. The flow decreases according to the specified pump coast down curve,
and the ASVs and NCVs open approximately 5 seconds later. RELAP is a very
versatile code in this way, though it does have some limitations that were observed
as described in this section.
Table 2.1 lists some geometric parameters for the core
exterior that are used in the RELAP inputs and are the same for all MITR-II fuel
cases studied.
2.4.1
Time Step Configuration
A time step sensitivity study was conducted in order to select a reasonable time step
size (to minimize RELAP code runtime) while preserving time resolution and accuracy for the LOF transient simulation. The reference RELAP input for the MITR-II
had a conservative time step configuration as shown in Table 2.2. The step size was
increased until the transient temperature output was noticeably different and the so-
38
Flow area (m 2 )
Volume (m 3 )
De (m)
Flow shroud
0.130
0.099
0.387
Mixing area
0.923
2.095
1.084
Hot leg
0.032
0.427
0.203
Cold leg
Downcomer 1
Downcomer 2
0.032
0.468
0.203
0.339
0.413
0.180
Downcomer 3
0.111
0.1256
0.076
0.016
Downcomer 4
0.029
0.018
0.063
0.22
0.04
4 NCVs
0.029
-
-
2 ACVs
0.007674
-
-
Table 2.1: Summary of primary loop dimensions in RELAP input. [13]
Region
lution remained stable.
The RELAP5 MOD3.3 time step input card takes the end time, minimum time
step, maximum time step and various other control options. When a run is initiated,
the step size starts at the minimum (1 ns), increasing by a factor of 1.1 with each
time step if the mass error is negligible, until the maximum time step size is reached.
The code terminates when it reaches the specified end time. If necessary, multiple
consecutive time periods can be specified to have different minimum and maximum
time step sizes.
The optimal time step configuration was found to be as shown in Table 2.3. Together the steady-state and transient cases run 85% faster with the new time-step
configuration. The optimized time step configuration was subsequently used for all
the fuel geometries discussed in this paper.
2.4.2
Limitations
While RELAP5 MOD3.3 is a versatile and powerful tool for thermal-hydraulic analysis, the user must be aware of how the results are calculated in order to identify
39
Table 2.2: Reference time step configuration [2]
Condition
Steady-state
Transient
Condition
Steady-state
Transient
Start times (s)
End times (S)
Min step size (s)
Max step size (s)
0.0
100.0
1 x 10-9
2 x 10-3
100.0
100.7
100.7
102.0
1 x 10-9
1 x 10-9
5 x 10-5
4 x 10-6
102.0
120.0
1 x 10-9
2.5 x 10-5
120.0
200.0
1 x 10-9
5 x 10-5
Table 2.3: Optimal time step configuration
Start times (s) End times (S) Min step size (s)
Max step size (s)
0.0
100.0
1 x 10-8
0.01
100.0
100.7
1 x 10-9
1 x 10-4
100.7
102.0
1 x 10-9
1 x 10-5
102.0
200.0
1 x 10-9
1 x 10-3
any inaccuracies or unphysical behavior. Some difficulties with RELAP5 MOD3.3 including inappropriate heat transfer model usage and boiling failure were experienced
during this work and are illustrated and explained in this section. Some of these
issues may be resolved by using a newer version of RELAP. The steady-state results
have been checked against hand calculations, though the transient results should be
interpreted with caution due to the findings discussed in this section.
Heat Transfer Modes
A study by K. Shirvan identified that the standard RELAP5 MOD3.3 heat transfer
model does not account for laminar flow, which contradicts the manual. [23] Instead
of using the Churchill-Chu correlation in the laminar flow region, RELAP5 MOD3.3
extrapolates the heat transfer coefficient down with the same slope. This result was
confirmed as shown in Figure 2-8 where the heat transfer coefficient as a function of
the Reynolds number linearly continues down into the laminar flow region, when a
constant heat transfer coefficient for laminar flow between parallel plates of MITR-II
fuel geometry (dashed line) is expected. This is of particular interest because after a
40
LOF the Reynolds number, corresponding to the natural circulation flow rate through
the coolant channels, is in the laminar flow region (indicated by the solid vertical line
in Figure 2-8). As can be seen in Figure 2-8, RELAP possibly over-estimates the heat
transfer coefficient and therefore transient temperatures may be higher than predicted
by RELAP.
RELAPS MOD3.3 data
-'15
10
Laminar flow-
region
U
Laminar flow between parallel plates
5000
Re# for natural
10 000
Reynolds Number
15 000
convection in MITR
Figure 2-8: RELAP5 MOD3.3 treatment of heat transfer in the laminar region. The
solid blue line represents RELAP data, the dashed line represents the constant heat
transfer coefficient for laminar flow between parallel plates of MITR-II fuel geometry.
[12] The laminar flow region is shaded in gray and the Reynolds number corresponding
to the natural circulation flow rate is indicated by the vertical solid line.
Boiling Analysis
While RELAP5 MOD3.3 provided for analysis of subcooled conditions, it was not
capable of evaluating boiling conditions beyond ONB. This was partially due to a
design fault in the simulated fuel models due to the hot channels. The hot channel
boils first, and by the time any other channels experience boiling, the hot channel
has experienced significant voiding, at which point the simulation aborts. Because of
41
this premature termination upon boiling in the hot channel, boiling in other channels
besides the hot channel cannot be observed. This can be overcome by removing the
hot channel.
In a 1997 study comparing RELAP5 MOD3.3 with other thermal-hydraulics codes
and experimental data, it was concluded that RELAP5 MOD3.3 use should be limited
to transients that do not result in a significant amount of boiling and voiding. [32]
One of the objectives of this study was to identify ONB in the hot channel, which was
found by observing an abrupt change in the heat transfer coefficient at the hottest
axial node in the hot channel. Because this was the first point of boiling in the en-
tire core, RELAP5 MOD3.3 ran the simulations without trouble up to power levels
slightly beyond the ONB-limiting power level. Phenomena such as OSV and OFI
were less easily observed because the power level at which these phenomena occurred
was typically the power level at which RELAP returned a high temperature or void
fraction error and terminated the simulation.
42
2.5
Previous LEU Fuel Design
A report produced by ANL in July 2013 presents a candidate fuel design that uses
18 finned fuel plates, thinner cladding, thinner coolant channels, and high density
U-Mo monolithic alloy fuel enriched up to 19.75%
23 5U.
[9] Transient analysis studies
for reactivity insertion accident (RIA) and loss of flow (LOF) scenarios showed that
the maximum fuel temperature would not reach the softening temperature of the aluminum cladding (approximately 450 C). More recent studies by ANL suggest several
designs with 17-19 unfinned plates, various fuel-to-clad thickness ratios, minimum
required power uprate to 6.7 - 7.1 MW to maintain flux, and ONB-limiting power
levels of 8.6 - 10.0 MW. [2]
The latest candidate fuel designs have graded fuel meat thickness, with outer fuel
plates containing thinner but uniform fuel meat.
While the fuel meat thickness is
graded, the overall plate thickness and coolant channel width are the same. Varied
fuel meat thickness makes for simpler geometry and improved manufacturability than
using finned fuel plates or increasing the coolant channel gap width next to hotter
plates in order to reduce power peaking.
Currently the hot channel (most limiting) in the MITR-II core is the end channel
(for even fuel meat thickness across plates) due to power peaking caused by extra
moderation of the end fuel plate when in the C-ring. In order to reduce this power
peaking in the end channel, the outer 3 fuel plates will have graded, thinner fuel meats
than the inner plates, creating a more even temperature profile across the fuel element,
reducing the peak-to-average fuel temperature ratio and increasing the limiting power
level for the core. This in turn makes it more difficult to identify the hot channel,
so each type of channel must be studied to ensure identification of the hot channel. [2]
43
2.6
High-Density Monolithic LEU U-10Mo Fuel
A high-density monolithic LEU U-1OMo fuel has been under development at ANL
since 1996 to accommodate conversion requirements of HEU-fueled research reactors
such as the MITR-II. The U-1OMo fuel can be manufactured as a dispersion-type fuel
and as a monolithic composition. The dispersion-type is of lower density and would
not provide sufficient excess reactivity in the MITR core geometry. The monolithic
fuel is therefore selected as the fuel meat for the MITR conversion. The constituents
of the LEU U-lOMo monolithic fuel used in this study are listed in Table 2.6.
Figure 2-9 is a schematic of an LEU monolithic U-10 Mo fuel plate. A zirconium
foil is placed between the U-10 Mo fuel meat and the Al-6061 clad, as shown in Figure 2-10, to facilitate interfacing and to reduce internal stresses caused by irradiation.
[21, 8] All the following fuel types discussed have the fuel meat composition listed in
Table 2.6 and fuel plate configuration as shown in Figure 2-9.
2.6.1
Material Properties
Tables 2.5, 2.6 and 2.7 list the thermal properties of aluminum-6061, zirconium and
U-10Mo, respectively, used in the thermal-hydraulic analyses. The RELAP5 MOD3.3
inputs take material properties in table form, where the tabulated values are obtained
from empirically-derived equations.
[10] To model finned fuel, the aluminum prop-
erties were multiplied by a fin factor of 0.743. The fin factor is used to account for
the difference in heat transfer ability when the surface is finned. While the fin factor
reduces the magnitude of the thermal conductivity and volumetric heat capacity, the
surface area is greater with fins, thus the heat transfer is slightly improved.
2.6.2
Fuel Temperature Limit
A downside to using U-1OMo fuel is that the temperature limit is lower than that
for UAl
cermet fuel. The melting point of UAl
44
is 1400 C while that of U-1OMo
Al-6061 Clad
mAI
U-10Mo
ad
U-Iassasi
O
Zr foi 1
aso
Zr foil
Figure 2-9: Fuel plate layer configuration. (Not to scale) [6]
Figure 2-10: Optical microscopy images of the U-10Mo fuel-Zr diffusion barrier inter-
face. [4]
Table 2.4: Composition of LEU U-10Mo monolithic fuel. [2]
Atomic Density (atoms/barn-cm)
Density (g/cm 3
)
Isotope
92
Mo
1.578 x 10-
3
0.2408
94
Mo
Mo
9.857 x 10-4
1.699 x 10-3
0.1537
95
96
Mo
1.781 x 10-
97
Mo
98
Mo
1.021 x 10- 3
2.584 x 10-3
1.033 x 10-3
1.025 x 10-4
7.751 x 10-3
1.798 x 10- 4
0.07046
238U
3.082 x 10-2
12.18
Total
4.953 x 10-2
17.02
00Mo
234U
235U
236U
45
0.2677
0.2837
0.1644
0.4202
0.1713
0.03983
3.025
Table 2.5: Thermal properties for Al-6061. [10]
Temperature ( C) Thermal Conductivity (W/mK) Vol. Heat Capacity (J/m3 K)
20
260
537.8
1648.9
167.5
184.3
201.1
268.1
2.539
2.688
2.828
3.391
x
x
x
x
106
106
106
106
Table 2.6: Thermal properties for zirconium. [10]
Temperature ( C) Thermal Conductivity (W/mK) Vol. Heat Capacity (J/m 3 K)
19.95
204.4
19.08
277.4
18.96
371.1
19.11
537.8
-
-
93.3
1.866 x 106
2.019 x 106
-
21.35
-
20
2.244 x 106
Table 2.7: Thermal properties for LEU U-Mo. [10]
Temperature ( 0 C)
Therma Conductivity (W/mK)
(unirradiated/irradiated)
Vol. Heat Capacity (J/m3 K)
10.64/9.30
2.362 x 106
93.3
-
2.430 x 106
204.4
-
2.538 x 106
315.6
-
2.655 x 106
426.7
-
2.781 x 106
800.0
37.36/32.67
46
-
20.0
is 1135 C. [13] Recent U-1OMo mini-plate irradiation testing has shown that blistering occurs at temperatures as low as 365 C, which is lower than the aluminum
softening temperature of 450 C. [24, 4]. Since conclusive results on the U-1OMo fuel
temperature limit is not yet available, it is conservative to assume that the peak fuel
temperature limit is 365 C for the purposes of this study.
2.6.3
Graded Monolithic LEU U-10Mo Fuel Meat
The current MITR-II power level is limited by power peaking in the end channels.
In order to flatten the power profile, an RERTR study proposed a test matrix of
graded fuel elements. Graded in this context means progressively thinner fuel meats
towards the end plates. Fuel designs with 12 or 15 mil unfinned cladding thicknesses,
17-19 plates and fuel meat gradation combinations as listed in Table 2.8 were tested.
STAT7 3 , MCODE and MCNP5 were used to analyze neutronic (cycle length, power
profiles etc.) and thermal-hydraulic (temperatures, ONB margin etc.) performance
of the fuel types. [2] All the fuels in this test matrix were unfinned. Unfinned fuel
accommodates thinner clad and therefore the fuel can be distributed among more fuel
plates per element, reducing the power density per plate. Furthermore, the milling
of very precise fins onto each fuel plate is a time-consuming and costly process. A
fuel design without fins is desirable for these thermal-hydraulic and economic reasons.
Each case was labeled to indicate the number of plates, fuel meat thickness grading configuration and nominal fuel meat thickness. For example, 19B25 (the most
promising combination, and the combination analyzed in this study) has 19 fuel
plates, grading configuration B and 25 mil nominal (interior plate) fuel thickness.
Figure 2-11 shows that power peaking in the end channel is significantly reduced by
grading the fuel meat thickness such that the outer plates contain less fuel.
3
STAT7 is a new computer code under development at Argonne National Laboratory for steadystate, thermal hydraulic uncertainties analysis of plate type fuel reactors such as the MITR.
47
80
8o
-constant
70
-
i
reduced
meat thickness
meat thickness - Combination
70
A
60
60
50
50
40
z
40
30
30
fat
20
20
.
N
X
U.
S
10
0
1
2
3
5
4
6
7
8
10
9
11
12
13
14
15
16
17
18
80
-constant
meat thickness
-reduced meat thickness - Combination
70
8
60 S
N
50
40
30
a
a:
20
Z
10
hIIIIIII
1
N
2
3
5
4
6
7
8
9
10
k0
11
-constant
12
13
14
15
16
17
18
meat thickness
-reduced meat thickness - Combination C
70
60
50
x
SD
D
40
iSt
30
20
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
80
.0
--- constant meat thickness
70
-reduced meat thickness - Combination
- mee70
a
60
60
ESD50
so
40
40
U 30
30
20
20
N
a
10
0
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
#
Plate
Figure 2-11: End plate heat flux reduction caused by graded fuel meat thickness for
gradation combinations A, B, C and D as listed in Table 2.8. [2]
48
Table 2.8: Graded fuel meat thickness combinations. [2]
Combination
Fraction of nominal meat thickness (%)
Interior plates 2nd plate 3rd plate 4th plate End plate
A
100
100
B
100
C
D
100
100
60
45
100
60
70
70
55
70
80
70
80
50
60
50
60
49
50
Chapter 3
Research Objectives
The objective of this study was to perform steady-state and transient analyses (for
a LOF accident scenario) for the proposed 19B25 unfinned graded monolithic LEU
U-10Mo MITR fuel design using RELAP5 MOD3.3. A preliminary objective was to
observe the effect of fin removal on the steady-state and transient temperature and
flow behavior compared to a finned reference case.
First the steady-state and transient cases for an unfinned 18-plate monolithic LEU
U-10Mo fuel type was compared to that of an 18-plate finned monolithic LEU U-10Mo
fuel (resembling current MITR-II fuel but with 3 extra plates and LEU fuel meat).
This comparison would show that the fuel needed re-designing so a high enough power
level could be achieved to compensate for the loss in neutron flux due to parasitic
absorption in
238
U after LEU conversion.
The second objective of this study was
to perform steady-state and LOF transient analyses for the proposed 19B25 graded
monolithic LEU U-10Mo fuel (see Section 5) for the proposed power level of 7.0 MW
for beginning-of-life (BOL), middle-of-life (MOL) and end-of-life (EOL) cores. Oxide
and thermal conductivity sensitivity studies were also performed to characterize the
effect of burnup on thermal properties of the fuel.
It was expected that unfinned fuel elements would operate at a higher temperature because of the decreased surface area compared to the finned plate design. The
51
hot channel temperature could be reduced and the radial power profile across each
element could be flattened by reducing the fuel meat thickness in the outer six plates
(three on each end) of each element.
Analyses were performed using a RELAP5
MOD3.3 model of the proposed 19B25 fuel type which was checked against a semianalytical 2D thermal-hydraulic model built in Mathematica. The overall objective
was to observe temperature, flow and boiling behavior during a LOF accident tran-
sient to determine if the 19B25 fuel is suitable for the MITR-II LEU conversion. In
order to qualify, the 19B25 fuel channels must not experience ONB at less than 9.8
MW (40% margin to the proposed uprate power of 7.0 MW), the maximum clad
temperature must not exceed the clad softening temperature of 450 C, and the fuel
must not exceed the U-1OMo blistering temperature of 365 C.
52
Chapter 4
Effect of Fin Removal
The current MITR-II fuel plates are milled with tiny fins for cooling. While the fins
increase the surface area to improve heat transfer, they limit the fuel redesign in several geometric and thermal-hydraulic ways. With the proposed power uprate, more
plates per assembly will be necessary in order to reduce the power density per plate.
The fins effectively double the minimum clad thickness because the clad touching the
fuel needs to be a certain thickness to keep fission products in and the fins cannot be
accurately milled within tolerances below a certain thickness.
Babcock & Wilcox, the manufacturer of MITR-II fuel, currently uses a custom
process to mill the fins with precision that is slow and costly. If the clad thickness was
reduced to accommodate more plates per element, it may not be possible to fin the
clad without breaching through to the fuel. Removing the fins allows the plates to be
spaced closer and have thinner effective cladding thickness (therefore making room
for more plates per assembly), and saves time and money. This chapter explores the
effects of fin removal from a 19-plate LEU fuel design.
53
4.1
MIT27 Reference Case
The reference case for this study is an 18-plate, finned, high-density monolithic LEU
U-1OMo fuel with the same fuel meat thickness in every plate, which is referred to as
MIT27. This fuel type is essentially the same as the currently-used HEU fuel type,
except for the number of plates per element (18 instead of 15), the addition of a zirconium foil between the fuel and clad, and the fuel meat (high density monolithic LEU
U-1OMo instead of HEU U-AlX). MIT27 was chosen to be the reference case because
it was the benchmarked reference case for the ANL candidate fuel development study.
[2]
Figure 4-1 shows the nodalization diagram for the MIT27 case. RELAP5 models
the entire primary loop (with temperatures constrained at the inlet and outlet of
the heat exchanger), but the structures of interest are 302 (average inner channel),
312 (average end channel), 402 (hot inner channel) and 412 (hot end channel). The
average inner (not end) coolant channels are lumped together into pipe 302 with a
corresponding heat structure labeled 1302.
All the heat generated in the lumped
1392 heat structure is deposited in the coolant flowing through pipe 302. The fins are
not geometrically defined as fin structures (RELAP doesn't accommodate geometric
specifics) and so the fins are "homogenized" by conversion factors to account for the
"fin effect". [2, 9] The effective channel gap size and other geometric quantities that
account for the fins are listed in Table 4.1.
In order to analyze each unique type of channel and plate (average/hot, inner/end), the plates are split in half as shown in Figure 4-2. The hot channel/plate
analyses are performed on hot "stripes"
(1/4
channels and 1/4 half plates) rather than
on entire channels/plates, as shown in Figure 4-2 (though the geometry and power
profiles provided in this section are for whole channels and half-plates). The limiting
stripe in this core is the hot end stripe (412). Table 4.1 lists all the dimensions for
each type of channel. The word "channel" is used to describe both a physical coolant
54
103
Mixing area 2
uppplin snglvol
100
Hot log
-
10 1
snkref
tmdpvol
102
Cold leg
,
cidleg
tmdpvol
201
Pump
tmdpjun\
(trip 403)
105
Mixing area 1
uppl2 snglvol
202
ASV valve
trpvlv
(trip 401)
N
203
Downcomer 1
regnl
pipe
109
Mixing area 3
uppl4
snglvol
108
Flow shroud
uppl3 snglvol
0
----------------------- -- -------0
LO
I-
I
205
Downcomer 2
regn2
pipe
208
NC valve
trpvlv
(trip 402)
0
w
__
0
0
N
A
-CD--u
(0NA
0_
.
1
j
1
~(50
--
0
lD
0
207
owncomer 3
regn3
pipe
1
210
Downcomer 4
regn4
pipe
110
Fuel bottom
inltpl snglvol
211
Figure 4-1: RELAP model of the MIT27 fuel case.
There is a separate RELAP
channel for each unique coolant channel and plate combination.
55
Table 4.1: MIT27 F uel Geometry
Geometry
Dimensions
No. fuel elements
22
18
No. plates per element
Plate length
Plate width
Fuel width
23"
0.5842 m
2.308"
0.05862 m
0.05288 m
2.082"
No. ave. interior channels per element
No. end channels per element
17
2
No. hot interior channels per core
1/s
No. hot end channels per core
Ave. interior channel flow area
Ave. end channel flow area
1/4
0.1892
0.0992
0.1629
0.0813
Hot interior channel flow area*
Hot end channel flow area*
Ave. interior Dh
sq.in
sq.in
sq.in
sq.in
0.0806"
0.0426"
0.0695"
0.0350"
0.009016"
Ave. end Dh
Hot interior Dh
Hot end Dh
U-1OMo fuel meat thickness
Al clad thickness
Zr foil thickness
Eff. interior hot ch. oxide thickness
Eff. end hot ch. oxide thickness
Fin height/width/gap
Ave. interior channel eff. gap width
Ave. end channel eff. gap width
Hot interior channel gap width
Hot end channel gap width
Heat transfer surface area per plate
0.015"
0.001"
0.008"
0.006"
0.01"
0.082"
0.043"
0.0706"
0.0352"
5.32 sq.in
1.22 x 10-4 m 2
6.40 x 10-5 m 2
1.05 x 10-4 m 2
5.24 x 10-5 m 2
2.05 x 10-3 m
1.08 x 10-3 m
1.77
8.88
2.29
3.81
2.54
2.03
1.52
2.54
m
m
m
m
m
m
m
m
2.08 x 10-3 m
1.09 x 10-3 m
1.79 x 10-3 m
8.94 x 10-4 m
3.43 x 10-3 m2
* All dimensions are given for a whole channel.
56
x 10-3
x 10-4
x 10-4
x 10-4
x 10-5
x 10-4
x 10-4
x 10--4
I
channel between plates and a lumped pipe/channel in RELAP, though care is taken
to clarify which type is implied.
The power profiles for each heat structure were generated using neutronics code
MCODE and parsed into RELAP in 18 axial nodes per plate type. Figure 4-3 shows
the power profiles for each type of plate (half plate) in the MIT27 fuel RELAP model.
4.1.1
Steady-State Analysis
Steady-state analysis of the MIT27 reference case shows that the hot end channel
(most limiting channel) wall temperature exceeds the saturation temperature at 7.0
MW, as shown in Figure 4-4. This means that the MITR fueled with MIT27-type fuel
would not be permitted to operate at 7.0 MW. A margin of at least 20% is required
between the licensed power level and the ONB-limiting power level.
Ave. Inner Channel (302)
Ave. End Channel (312)
Hot Inner Channel (402)
Hot
End
Channel
(412)
i
{
20
K
ti4
V
,
15
/
'II
/
010
I
i
5
2
r
r
t
0
's
60
80
100 120 140
60
80
100 120 140
60 80
Temperature (*C)
100 120 140
60
80
100
120
140
Figure 4-4: Axial temperature profiles for MIT27 fuel (finned) for steady-state operation at 7.0 MW
57
412 Hot X End
Channel Strips
-312
- -
-
1412
Hot'%
End
-t - --42
/
__
1 02 A-rage banrs - 4lat-
F Plate Stripe
1402 Hot %
Inner
-a
-
% Inner
Channel Stripe
402 Hot
Figure 4-2: Plate and channel stripe configuration illustration for the MIT27 fuel
case.
axial 1
axial 2
axial 3
axial 4
axial 5
axial 6
axial 7
axial 8
axial 9
axial 10
axial 11
axial 12
axial
axial
axial
axial
axial
axial
13
14
15
16
17
18
1302
27.8
24.5
28.6
33.1:
37,4
41.3
45.1
48.5
51.3
62.8
59.0
62.7
63.7
62.8
60.9
57.1
52.8
64.2
1312
27.8
24.5
28.6
33.1
37.4
41.4
45.2.
48.5
51.3
62.9
59.1
62.7
63.7
62.9
60.9
57.1
52.8
64.2
1402
1412
23.3
28.8
33.9
39.8
49.5
61.2
74.1
94.1
116.7
131.1
136.4
139.2
139.7
136.8
132.5
141.6
23.3
28.8
33.9
39.8
49.5
61.2
74.1
94.1
116.7
131.1
136.4
139.2
139.7
136.8
132.5
141.6
Figure 4-3: Power profiles for each unique type of half-plate in the MIT27 fuel, where
1302 is the average inner half plate, 1312 is the average end half-plate, 1402 is the
hot inner half plate, 1412 is the hot end half plate.
58
4.1.2
Loss-of-Flow
During a LOF accident after steady-state operation at 7.0 MW fuel type MIT27 does
not perform well. Figure 4-5 shows that the most limiting hot stripe wall temperature
exceeds the saturation temperature, and that there are temperature oscillations, flow
oscillations and void fraction deviations from 0.0. Upon LOF, the hot end channel
wall temperature rises above the saturation temperature and holds steady due to
boiling. When the wall temperature exceeds the saturation temperature, water on
the surface starts to boil (ONB), the heat transfer coefficient increases, more heat is
removed by the coolant, and the wall temperature cools down below the saturation
temperature and the boiling stops. This process repeats within a short period of time
to create the oscillatory and OFI behavior observed in Figure 4-5. This process of
periodic boiling during a temperature transient is not necessarily a dangerous condition as long as the oscillation dampens over time. Once ONB is surpassed by OFI,
acceleration to CHF may become a concern.
Though the maximum fuel and clad
temperatures do not approach the softening temperature of Al-6061, the MITR-II
should not nominally operate at 7.0 MW with this type of fuel if nucleate boiling is
present during steady-state operation and OFI is a consequent effect of a LOF acci-
dent. This illustrates the need for fuel geometry redesign to accommodate the LEU
fuel.
4.2
MIT30 - The Effect of Removing Fins
Fuel type MIT30 was studied as an intermediate case to characterize the steady-state
effect of changing finned fuel plates to unfinned (smooth).
Without fins the heat
transfer surface area is reduced without changing the effective coolant flow area, as
shown in Figure 4-6. Table 4.3 lists the geometric factors for the unfinned fuel. Removal of the fins causes an increase in the fuel and clad temperatures throughout the
fuel element (see Figure 4-7) and causes a significant reduction in the margin to ONB
(see Table 4.2). The ONB-limiting power level was reduced by a factor of more than
59
160
Hot End Channel (412) At Hottest Axial Point
-
140
Channel wall temp.
Saturation temp.
-
Bulk temp.
0
cd
Max. fuel temp.
-
120
.
,.
ci.
Su 100
80
-20
0
20
40
60
Time after SCRAM (s)
80
100
Hot End Channel (412)
0.020
0.4
0.015
0
0.2
0
0.010
I
__
void fraction
0.
:3
0.005
Ann
mass flow
-*UU120
0
20
40
60
80
100
Time after SCRAM (s)
Figure 4-5: MIT27 LOF accident hot channel temperatures (top), and void fraction
and flow rate (bottom) after steady-state operation at 7.0 MW.
60
two, well below the desired uprate power level of 7.0 MW. The position of the peak
axial point was expected to remain the same with the change to unfinned fuel plates
because the power profiles for each type of plate were kept the same. Figure 4-8 shows
ONB by the discontinuity in the otherwise linear plot of heat transfer coefficient as a
function of steady-state operating power level. No transient analyses were performed
for the MIT30 case as it is clear from the steady-state results that boiling phenomena
beyond ONB will occur during LOF, and this fuel type is not a feasible candidate for
the conversion.
Figure 4-6: The effective coolant channel gap width remains the same after removal
of fins from MIT27 to create fuel MIT30.
Table 4.2: ONB-limiting power level with core mass flow of 1800 gpm (112 kg/s).
Limiting ONB
Power Level (MW)
MIT27 (finned LEU)
MIT30 (unfinned LEU)
5.5 MW
2.5 MW
61
Position
Hot stripe 402, Axial node 7
Hot stripe 412, Axial node 7
Ave. Inner Channel (302)
2
0
0
Ave. End Channel (312)
Hot Inner Channel (402)
Hot End Channel (412)
0
15
0
110
/
0
5
i
1\
0Q
60
80
100 120 140
60
80 100 120 140
60
80
100 120 140
60
80
100 120 140
Temperature ('C)
Figure 4-7: Increase in wall temperature at 7.0 MW when fins are removed. Finned
(MIT27) and unfinned (MIT30) channel wall temperatures are represented by dashed
and solid lines respectively. The maximum wall temperature increase is approximately
30 C.
30
'7
25
o 20
0
0
finnedd
i
15
x
10
0
2
4
Steady-state power level (MW)
6
8
Figure 4-8: ONB determined by discontinuity in heat transfer coefficient vs.
power
curve. The power level at which the discontinuity occurs is the ONB-limiting
power
level.
62
Table 4.3: MIT30 Fuel Geometry
Geometry
Dimensions
No. fuel elements
No. plates per element
22
18
Plate length
Plate width
Fuel width
23"
0.5842 m
2.308"
0.05862 m
2.082"
0.05288 m
No. ave. interior channels per element
No. end channels per element
17
2
No. hot interior channels per core
No. hot end channels per core
1/8
1/4
Ave. interior channel flow area
Ave. end channel flow area
0.1892 sq.in
0.0992 sq.in
0.1722 sq.in
0.0860 sq.in
0.1584"
Hot interior channel flow area
Hot end channel flow area
Ave. interior Dh
Ave. end Dh
0.0844"
Hot interior Dh
0.1445"
0.0733"
0.009016"
Hot end Dh
U-1OMo fuel meat thickness
Al clad thickness
Zr foil thickness
0.015"
Interior hot channel oxide thickness
End hot channel oxide thickness
0.004"
0.003"
0.082"
0.043"
0.001"
Ave. interior channel gap width
Ave. end channel gap width
Hot interior channel gap width
Hot end channel gap width
Heat transfer surface area per plate
63
1.22 x 10-4 m2
6.40 x 10-5 m2
1.11 x 10-4 m2
5.54 x 10-5 m2
4.02 x 10-3 m
2.14 x 10-3 m
3.67 x 10-3
1.86 x 10-3
2.29 x 10-4
3.81 x 10-4
2.54 x 10-5
2.03 x 10-4
1.52 x 10-4
m
m
m
m
m
m
m
2.08 x 10-3 m
1.09 x 10-3 m
0.0746"
0.0372"
1.90 x 10-3 m
9.45 x 10-4 m
2.66 sq.in
1.72 x 10-3 m 2
64
Chapter 5
Proposed New Fuel Design
The candidate fuel type selected for the MITR-II conversion is an unfinned 19-plate
fuel with 12-mil Al-6061 clad and LEU monolithic U-10Mo fuel meat (permutation
19B25 in the RERTR LEU core design paper). [30] The fuel meat thickness in each
fuel plate is graded according to configuration B (see Section 2.8) where the outer
six plates (3 on each side) have progressively thinner fuel meat towards to outside
of the assembly to reduce power peaking due to extra moderation at the assembly
perimeter. The fuel meat thickness grading is listed in Table 5.1 along with other
fuel geometry dimensions for all three cases studied.
With the graded fuel meats, there are a greater number of unique channels, plates
and interfaces. In order to properly identify the hot channel, and characterize the
thermal-hydraulic behavior of each type of channel during a LOF, each type of channel (including average and hot channels) was defined in the RELAP input. Figure 5-1
shows the channel and plate numbering convention, and Figure 5-2 shows the com-
plete core and primary coolant system nodalization diagram for the 19B25 RELAP
model.
A hot plate produces more power than surrounding plates and therefore depletes
faster.
After some time, the hot plate has depleted more than other plates to the
point where it is no longer the hottest and a different plate becomes the hot plate.
65
Thus three 19B25-fueled cores were studied with varied burnup.
F
xa
x2
-
<
x2
W0x
E
x2
N
~
x2
xio
x12
beginning with the number 3 or 13 indicate average channels or plates respectively (as
shown in the figure). Labels beginning with the number 4 or 14 indicate hot channels
or plates respectively (not shown).
5.1
Fuel Cycle and Core Power Distributions
The proposed conversion plan is to replace the last HEU core with an entirely fresh
LEU core of fuel type 19B25 with only 22 elements. Nominally refueling operations
are performed every 3-4 months where 3 new elements are introduced, 3 old are removed to the wet storage ring, and the remaining elements are shuffled. After the
fresh LEU core has been irradiated for several months, the normal refueling procedure will be resumed to use up to 25 of the 27 fuel element spaces (2-3 are used
for in-core experiment). In the fresh core case the power is concentrated in a fewer
number of elements (to limit fresh core reactivity in order to maintain an adequate
shutdown reactivity margin) and in the bottom half of the core (the control blades
start low), making it the most conservative case. This configuration is referred to as
beginning-of-life (BOL), as opposed to beginning-of-cycle (BOC) for a core made up
of a mixture of fresh and partially depleted fuel elements.
66
103
Mixing area 2
upppln snglvol
100
Hotreog
snkpref
101-
102
o
__
__
-
~
-
coldleg
-;
\ tmdpvol
105
Pump
Mxnara1tdjn
____
___202
1
0V
I
_
j-
108
Flow shroud
203
-
ASV volve
trpvlv
(trip 401)
_
Downcomer 1
..
regnl
pipe
-
109
Mixing area 3
0o
-
upp13 snglvoluppl4
uppl
sngvolsnglvol
-
W
W
II
IW
205
Downcomer 2
17
w j
W
l Wi
W
1
-
(trip 403)
-
_____
uppl2 snglvol j
re n2
0
U
_
I...
_
__w_--20
NC volve
-\X
(trip 402)
PQ
.9
N
D
w
3
D
<
N
3
D
'ao
aregn3
pipe
a(
-
I- I
ii
-A
1
o
pipe
110
Fuel bottom
211-
___
---------------
i-i
C L
CD
fl CD-CDDl
207
N
W
N
3
'r
4:
DN
<N.D:N
Downcomer
3
^
inltpl snglvol
Figure 5-2: RELAP model of the 191B25 fuel case. There is a separate RELAP channel
for each unique coolant channel and plate combination.
67
Table 5.1: 19B25 Fuel Geometry
Geometry
Dimensions
No. fuel elements
No. plates per element
Plate length
Plate width
Fuel width
23"
2.308"
2.082"
No. ave. interior channels per element
No. 4th channels per element
No. 3rd channels per element
No. 2nd channels per element
No. end channels per element
No. hot channels per core**
Ave. interior flow area
Ave. end flow area
Hot interior flow area
Hot end flow area
Ave. interior Pw
Ave. end Pw
Hot interior P
Hot end Pw
Interior Ph
End Ph
U-10Mo fuel meat thickness, inner
U-10Mo fuel meat thickness, intermediate
U-10Mo fuel meat thickness, end
Al
Al
Al
Zr
clad thickness, inner
clad thickness, intermediate
clad thickness, end
foil thickness
Interior channel gap width
End channel gap width
Ave. heat transfer surface area
Hot heat transfer surface area
22/24*
19
0.5842 m
0.05862 m
0.05288 m
12
2
2
2
2
5
0.1722 sq.in 1.11 x 10-4 r
0.1516 sq.in 9.78 x 10-5 r
0.1553 sq.in 1.00 x 10-4 r
0.1368 sq.in 8.82 x 10-5 r
4.7652"
0.1210 m
4.7474"
0.1206 m
4.3132"
0.1096 m
4.2954"
0.1091 m
4.164"
0.1058 m
2.082"
0.0529 m
0.025"
6.350 x 10-4 m
0.017"
4.318 x 10-4 m
0.013"
3.302 x 10-4 m
0.012"
3.305 x 10-4 m
0.016"
4.064 x 10-4 m
0.018"
4.572 x 10-4 m
0.001"
2.540 x 10-5 m
0.073"
1.854 x 10-3 m
0.039"
9.906 x 10-4 m
53.08 sq.in
0.0342 m 2
47.89 sq.in
0.0309 m 2
* The BOL core has 22 elements while the MOL and EOL cores have 24 elements.
** One for each unique type of channel.
68
Middle-of-life (MOL) and end-of-life (EOL) cores containing 24 fuel elements were
also analyzed for thermal-hydraulic performance during a LOF accident. The MOL
core represents the BOC of an equilibrium core with partly burned fuel but no xenon.
EOL represents the end-of-cycle (EOC) of an equilibrium core with deeply burned
fuel and xenon present. Figure 5-3 shows the shift in power distribution from BOL
through MOL to EOL. As the power peaking shifts from the A and B rings to the
C ring over the core life, the power distribution within each element also changes
slightly and was therefore analyzed in order to assure characterization of the most
limiting conditions the core sees throughout its life.
G=B
2 B312
2OL
104
CI'BUJ
B
(7
'
1
B97" A'B'A
(I
I2
MW
0.411
1
CIO(.0.
BOL
C1
MOL
-
3 C1+
- cl 0.243
- (2
EOL
Figure 5-3: Power distribution shift over core life, from BOL though MOL to EOL
at 7.0 MW. The gray positions represent unfueled positions.
Before the conversion can proceed, the high-density monolithic LEU U-lOMo fuel
must be approved by the manufacturer (Babcock & Wilcox) and qualified as suitable
for use in non-power reactors by the NRC. Though the manufacturer has demonstrated fabrication of monolithic LEU U-lOMo fuel test pieces, qualification is pending construction of the fabrication plant so that process-dependent results can be
approved by the NRC.
69
5.2
Axial Power Profiles
While neutronic analyses were not within the scope of this study, the power profiles
provided for the 19B25 fuel type were found using MCODE [33, 26] and parsed into
the RELAP model. RELAP needs the absolute power generation for each volume to
be specified so that it can calculate how much heat is deposited in the coolant and
other structural materials and at what rate. [26] Figures 5-4, 5-5 and 5-6 show the
power profile for a 19B25 fuel element at 3 stages, averaged over all fuel elements
(22 in the fresh core (BOL) and 24 in the MOL and EOL cores).' The BOL case
is most conservative because the excess reactivity is high and concentrated into 22
fresh elements instead of the nominal 24-25 elements in the MOL and EOL cores (2-3
positions are typically occupied by experiments or aluminum dummy elements), and
the control bank height starts relatively low so the power generated is further concentrated in the lower half of the core. The middle-of-life (MOL) core contains partially
depleted elements at the beginning of a cycle with no xenon present. The end-of-life
(EOL) core contains depleted elements at the end of a fuel cycle with equilibrium
xenon.
The power profiles for each of the average plates were made up of the average of
all the plates of that type from all the fuel elements in the core combined. As can be
seen in Figures 5-4, 5-5 and 5-6, of the average plates, the outer-most of the plates
with 100% fuel meat thickness (1312) are the hottest. On average the element-wide
power profile is relatively consistent. The hot plate does, however, change position
from BOL through EOL.
The hot plate power profiles for each unique type of plate was created by taking
the plate with the highest total power for that type of plate from the entire core of
22-24 MCODE-simulated elements. As shown in Figures 5-4, 5-5 and 5-6, the hot
plate shifted from the end channel (1442) at BOL to the 4th channel from the end
'The
color mapping is scaled and consistent across Figures 5-4, 5-5 and 5-6.
70
Plate 1
Axial1
Axial 2
Axial 3
Axial 4
Axial 5
Axial 6
Axial 7
Axial 8
Axial 9
Axial 10
Axial 11
Axial 12
Axial 13
Axial 14
Axial 15
Axial 16
Axial 17
Axial 18
Sum
I,
0.446
0,380
0,439
0.508
0.575
0.640
0.706
0.783
0.881
1.000
1.078
1.111
1.109
1.088
1.052
0.988
0.933
1.133
14.850
Plate 2 Plate 3 Plate 4 Plate S Plate 6 Plate 7 Plate 8 Plate 9 Plate 10 Plate 11 Plate 12 Plate 13 Plate 14 Plate 15 Plate 16 Plate 17 Plate 18 Plate 19
0.506
0.465
0.599
0.565
0.549
0.539 0.535 0.529 0.529 0.526 0.527 0.530 0538 0.554 0.586 0.456 0.501 0.450
0.437
0,401
0.521
0.493
0.477 0.467 0,462 0.459 0.456 0.457 0.458 0.462 0.471 0.487 0.518 0.401 0.442 0.393
0.507
0.468
0.613
0.580
0.563
0.551
0.547
0.543
0.541
0.540
0.543
0.546
0.555
0.573
0,608
0.468
0.512
0.452
0.587
0.542
0.708
0.672
0.651
0.641
0,632
0.629
0.628
0.626
0.631
0.637
0.647
0.668
0.708
0.544
0.595
0,525
0.663
0.616
0.805
0.763
0.741
0.727
0.719
0,714
0.711
0.713
0.715
0.723
0.736
0.759
0.801
0.617
0.676
0,594
0.740
0.683
0.895
0.848 0.825 0.809 0.799 0.796 0.792 0.794 0.797 0.805 0.820 0.843 0.891 0.684 0.748 0.658
0.815
0.753
0.984
0.936
0.907
0.890
0.880
0,875
0.872
0.872
0.876
0.883
0.898
0.925
0.977
0.752
0.823
0,724
0.897
0.822
1.075
1.016
0.985
0.965
0.955
0.948
0.944
0.945
0.949
0.959
0.977
1.005
1.064
0.819
0.897
0.794
0.995
0.902
1.170
1.102
1.067
1.043
1.033
1.023
1.022
1.021
1.028
1.037
1.053
1.088
1.156
0.892
0.983
0.875
1.108
0.993
1.272
1.193
1.152
1.125
1.109
1.099
1.098
1.099
1.107
1.121
1.145
1.188
1.270
0.996
1.110
1.011
1.183 1.058 1.351 1.262 1.213 1.185 1.169 1.161 1.155 1.158 1.167 1.186 1.213 1.263 1.361 1.076 1.217 1.129
1.212
1.083
1.378
1.289
1,238 1.207 1.189 1.182 1.180 1.183 1.190 1.209 1.238 1.292 1.392 1.104 1.255 1.173
1.211
1.079
1.373
1.282
1.231
1.202
1.184
1.176
1.173
1.175
1.183
1.202
1.235
1.289
1.392
1.104
1.259
1.178
1.186
1.055
1342 1.250 1.202 1.172 1.155 1.146 1.142 1.145 1.154 1.173 1.205 1.260 1.366 1.087 1.242 1.166
1.142
1.011
1.283
1.195 1.145 1.117 1.102 1.092 1.091 1.095 1.104 1.122 1.153 1.212 1.319 1.057 1.222 1.158
1.069
0.942
1,194
1.109
1.063
1.038
1.023
1.014
1.011
1.016
1.025
1.043
1.073
1.129
1.232
0.993
1.153
1.107
0.999
0.874
1.098
1.016
0.973
0.948
0.932
0.924
0.921
0.925
0.936
0.953
0.982
1.036
1.139
0.926
1.083
1.053
1.216
1.075
1.346
1.2;0
1.204
1.175
1.163
1.153
1.149
1.156
1.163
1.183
1.215
1.270
1.387
1.128
1.307
1.261
16.471
14.823
19.006
17.821
17.185
16.800
16.588
16.463
POWER PROFILE FOR EACH AVERAGE STRUCTURE
1302
1312
1322
1332
1342
Axial1
0.538
0.592
0.460
0.503
0.448
0.468
0.519
0.401
0.439
0.387
Axial 2
Axial 3
0.553
0.610
0.468
0.510
0.446
Axial 4
0.642
0.708
0.543
0.591
0,516
Axial5
0.729
0.803
0.616
0.670
0.584
Axial6
0.812
0.893
0.684
0.744
0.649
Axial7
0.892
0.981
0,752
0.819
0.715
Axial 8
0.968
1.070
0.820
0.897
0.789
Axial 9
1.047
1.163
0.897
0.989
0.878
Axial 10
1.130
1.271
0.995
1.109
1.005
Axial 11
1.194
1.356
1.067
1.200
1.103
Axial 12
1.218
1.385
1.093
1.233
1.142
Axial 13
1.212
1382
1.092
1.235
1.144
Axial 14
1.182
1.354
1.071
1.214
1.127
Axial 15
1.130
1.301
1.034
1.182
1.105
Axial 16
1.049
1.213
0.968
1.111
1.047
Axial 17
0.959
1.119
0.900
1.041
0.993
Axial 18
1.189
1.366 1.101
1,262
1.197
Sum
16.912 19.086 14.963 16.748 15.276
16.414
16.445
16.553
16.773
17.153
17.841
19.166
15.103
17.025
15.702
POWER PROFILE FOR EACH HOT STRUCTURE
1402
1412
1422
1432
1442
0.977
1.057
0.863
1.012
0.238
0.815
0.899
0.727
0.853
4.200
0.922
0.988
0.784
0.911
0.319
1.022
1.099
0.875
0.993
0.380
1.136
1.215
0.966
1.081
0.448
1,247
1.330
1.048
1.169
0.513
1.337
1.434
1.131 1.258 0.587
1.419
1.516
1.174
1.304
0.738
1.447
1.544
1.194
1.308
1.064
1.461
1.551
1.198
1.316
1.460
1.465
1.558
1.205
1315 1704
1.468
1.539
1.191
1.304
1.822
1.467
1.551
1190
1.305
1.834
1.411
1.500
1.153
1.272
1.811
1.344
1442
1.121
1.242
1.775
1.294
1.388
1.079
1.221
1.684
1.243
1.338
1.072
1.229
1.640
1.542
1.655
1.334
1.520
1.771
23.017 24.605 19.306 21.613 20.049
Figure 5-4: Beginning-of-life power profiles in kW per node volume, where the node volume is constant. Top: power profile
for each plate across the core-averaged fuel element. Bottom left: power profiles for each average heat structure for input into
RELAP, normalized for one plate each. Bottom right: power profiles for each hot heat structure for input into RELAP.
-
-
-
-
Table 5.2: BOL Hot channel factor, and power generation (% of total core power)
per plate type and per lumped heat structure for RELAP input.
Plate type
Power per plate Lumped Power Hot channel factor
Ave. inner 302
0.242
58.23
Ave. 4th
312
0.273
11.72
Ave. 3rd
322
0.214
9.192
Ave. 2nd
332
0.239
10.29
342
0.218
9.384
Hot inner
hot 4th
402
412
0.329
0.351
0.329
0.351
1.36
Hot 3rd
Hot 2nd
422
432
0.276
0.309
0.276
0.309
1.29
1.29
Hot End
442
0.286
0.286
1.31
Total
100%
72
-
Ave. End
1.29
Axial 1
Axial2
Axial 3
Axial4
Axial S
Axial 6
Axial 7
Axial8
Axial 9
Axial 10
Axial 11
Axial 12
Axial 13
Axial 14
Axial 15
Axial 16
Axial 17
Axial 18
Sum
Plate 1 Plate 2 Plate 3 Plate 4 Plate 5 Plate 6 Plate 7 Plate 8 Plate 9 Plate 10 Plate 11 Plate 12 Plate 13 Plate 14 Plate 15 Plate 16 Plate 17 Plate 18 Plate 19
0.413
0.477
0.443
0.575
0;549
0.535
0.526
0.522
0.520
0.520
0,520 0.523
0.528
0.537
0.553
0.585
0.455
0,497
0.441
0.349
0.407
0.378
0.496
0.472
0.460
0.452
0.447
0.445
0.446
0.447
0.450
0455
0.465
0.481
0.511
0.394
0.432
0.380
0.405
0.475
0.443
0.586
0.558
0,543
0.535
0,530
0.527
0.526
0.528
0.531
0,538
0;548
0.566
0.600
0.459
0.500
0.438
0,462
0.544
0.508
0,673 0.642 0.625 0:614 0.609 0.605 0.606 0.608 0.611 0,619 0.630 0.651 0.688 0,526 0.571 0.497
0.522
0,615
0.575
0.763
0.726
0.706
0.696
0.689
0.685
0.684
0.686
0.691
0.697
0.711
0.735
0.775
0.593
0.644
0.560
0.579
0.683
0.638
0.844
0.805
0.783
0.771
0.763
0.759
0.760
0.761
0.765
0.774
0.788
0.814
0.860
0.657
0.712
0.619
0.625
0.738
0.690
0.915
0.873
0.849
0.836
0.827
0.824
0.823
0.825
0.829
0.838
0.854
0.880
0.930
0.709
0.768
0.664
0,685
0.806
0.751
0.993
0.943
0.917
0.902
0.894
0.889
0.888
0.889
0.894
0.905
0.923
0.951
1.006
0.767
0.833
0.720
0.759
0.881
0.815
1,070
1.013
0.982
0.965
0.954
0.950
0.948
0.950
0.956
0.967
0.985
1.019
1.077
0.823
0.893
0.773
0.844
0,966 0.887
1.154
1.087
1.050 1.029 1.018 1.012 1.010 1.012 1.020 1.034 1.056 1.096 1.168 0.901 0.989 0.867
0.902
1.027
0.937
1.215
1.141
1.103
1.079
1.065
1.058
1,055
1.060
1.068
1.084
1.111
1.155
1.239
0.964
1.069
0.949
0.924
1.050
0.956
1.236
1.161
1.122
1.098
1.083
1.074
1.072
1.074
1.085
1.101
1.130
1.178
1.265
0.986
1.096
0.978
0.930
1.054
0.957
1.236
1.161
1.118
1.095
1.080
1.074
1.072
1.074
1.083
1.099
1.128
1.178
1.266
0.990
1.103
0.991
0.911
1.029
0.933
1.204
1.131
1.089
1.065
1.050
1.044
1.043
1.045
1.054
1.071
1.101
1.149
1.236
0.970
1.085
0.979
0.879
0.990
0.896
1.152
1.081
1.040
1.017
1.003
0.996
0.994
1.000
1.009
1.026
1.055
1.105
1.194
0,942
1.063
0.971
0.828
0.931
0.839
1.078
1.007
0.968
0.946
0.935
0.927
0.926
0.931
0.939
0.957
0.984
1.034
1.123
0.891
1.012
0.930
0,774
0.861
0.771
0,984
0.916
0.878
0.857
0.846
0.840
0.839
0.842
0.853
0.870
0,897
0.945
1.033
0.827
0.949
0.884
0.952
1.060
0.957
1.215
1.137
1.095
1.073
1.061
1.054
1.053
1.057
1.066
1.083
1.114
1.166
1.26Z
1.012
1.152
1.067
12,746 14.595 13.375 17.388 16.403 15.864 15.556 15.376 15.283 15.266 15.309 15.427 15.646 16.017 16.655 17.818 13.865 15.370 13.709
POWER PROFILE FOR EACH AVERAGE STRUCTURE
1302
-W
Axial 1
0,530
Axial 2
0.456
Axial 3
0,539
0.620
0.700
0.777
0.841
0.909
0.972
1.039
Axial 4
Axial 5
Axial 6
Axial 7
Axial8
Axial 9
Axial 10
Axial 11
Axial 12
Axial 13
Axial 14
Axial 15
Axial 16
Axial 17
Axial 18
Sum
1.089
1.107
1.106
1.077
1.030
0.959
0.871
1.087
15.709
1312
0.580
0.504
0.593
0.680
0.769
0.852
0,923
0.999
1.073
1.161
1.227
1.250
1.251
1,220
1.173
1.100
1.009
1.239
17.603
1322
1332
1342
0.449
0.386
0,487
0.427
0,420
0.451
0,488
0.558
0.630
0.697
0.753
0.819
0.887
0.978
1.048
1.073
0.365
0.422
0.480
0.541
0.517
0.584
0.647
0.700
0.759
0.819
0.894
0.950
0.971
0.973
1.078
0.952
1.057
0.919
1.027
0.865
0.972
0.799
0.905
0.985
1.106
13.620 14.982
0.599
0,645
0.703
0.766
0.856
0.926
0.951
0.960
0.945
0.925
0.879
0.829
1.009
13.227
POWER PROFILE FOR EACH HOT STRUCTURE
1402
1412
1422
1432
1442
0.957
1.044
0.846
0.982
0.877
0.802
0.877
0.709
0.825
0.750
0.892
0.966
0.767
0.872
0.802
0.996
1.072
0.841
0.942
0.871
1.101
1.184
0.924
1.036
0.960
1.204
1.295
1.007
1.120
1.040
1.282
1.357 1.057
1.185
1.093
1.337
1.426
1.104
1.224
1.134
1.436
1.351
1.103
1.215
1.084
1.360
1.442
1,102
1.199
1.062
1.374
1.452
1.111
1.202
1.057
1.369
1.446
1.102
1.203
1.052
1.349
1.421
1.087
1.189
1.054
1.314
1.388
1.059
1.160
1.025
1.243
1.318
1.022
1.118
0.994
1.183
1.268
0.983
1.097
0.975
1.132
1.229
0.966
1.101
1.003
1.416
1.531
1.223
1.388
1.269
21.664 23.153 18.013 20.055 18.100
Figure 5-5: Middle-of-life power profiles in kW per node volume, where the node volume is constant. Top: power profile for each
plate across the core-averaged fuel element. Bottom left: power profiles for each average heat structure for input into RELAP,
normalized for one plate each. Bottom right: power profiles for each hot heat structure for input into RELAP.
322
0.195
9.145
Ave. 2nd
332
0.214
10.06
-
Ave. End
342
0.189
8.881
Hot inner
hot 4th
402
412
0.309
0.331
0.309
0.331
1.38
1.32
Hot 3rd
Hot 2nd
422
432
0.257
0.287
0.257
0.287
1.32
1.34
Hot End
442
0.259
0.259
1.37
Total
100%
74
-
Ave. 3rd
-
-
-
Table 5.3: MOL Hot channel factor, and power generation (% of total core power)
per plate type and per lumped heat structure for RELAP input.
Plate type
Power per plate Lumped Power Hot channel factor
Ave. inner 302
0.224
59.02
Ave. 4th
312
0.251
11.82
Axial 1
Axial 2
Axial 3
Axial 4
Axial 5
Axial6
Axial 7
Axial 8
Axial 9
Axial 10
Axial 11
Axial 12
Axial 13
Axial 14
Axial 15
Axial 16
Axial 17
Axial 18
Sum
Cnf
Plate 1 Plate 2 Plate 3 Plate 4 Plate 5 Plate 6 Plate 7 Plate 8 Plate 9 Plate 10 Plate 11 Plate 12 Plate 13 Plate 14 Plate 15 Plate 16 Plate 17 Plate 18 Plate 19
0,455
0.531
0,494
0.643
0.611 0.594 0.584 0.577 0,572 0.570 0,568 0.569 0.572 0,580 0,596 0,628 0.486 0.532 0.471
0,403
0.468
0,434
0.569
0,539
0.522
0,511
0.505
0,500 0.499
0,499
0.502
0,509 0.519 0,537 0.574 0.446 0.496 0,447
0.501
0.576
0,530
0.693
0.654
0.633
0.620
0.612
0.607
0.607
0.606
0.610
0,620
0.635
0.663
0.712
0.559
0.631
0.577
0.593
0.678
0.622
0.808
0,761
0.735
0.719
0.709
0.705
0.705
0.706
0.711
0,724
0.743
0.779
0.844
0.668
0.759
0.699
0.676
0.770
0,704 0.914 0.860 0.829 0.812 0.801 0.795 0.793 0.797 0.804 0.817 0.843 0.883 0.959 0.761 0.867 0.801
0341
0.845
0.772
1.002
0.941
0.909
0.890
0.879
0.872
0.871
0.873
0.881
0.898
0.924
0.971
1.054
0.837
0.954
0.883
0,774
0.889
0.814
1.058
0.996
0.960
0.940
0,929
0.923
0.923
0.925
0.933
0.948
0.977
1.026
1.115
0.883
1.005
0.922
0.811
0.932
0.853
1.110
1.044
1.006
0.984
0.973
0.966
0.965
0.967
0.977
0.994
1.026
1.074
1.169
0.927
1.053
0.965
0.825
0.949
0.870
1.132
1.065
1.028
1.007
0.993
0.986
0.984
0.988
0.998
1.015
1.045
1.097
1.192
0.943
1.069
0.976
0.824
0.952
0.874
1.139
1.073
1.035
1.013
1.001
0.993
0.991
0.994
1.004
1.021
1.053
1.105
1.200
0.947
1.071
0.975
0.822
0.948
0.871
1.136
1.068
1.031
1.008
0.997
0.989
0.989
0,990
1.001
1.017
1.048
1.101
1.194
0.944
1.069
0.972
0.809
0.933
0.857
1.116
1.051
1.015
0.991
0.980
0.974
0.973
0.976
0.986
1.003
1.032
1.083
1.175
0.929
1.051
0.956
0.800
0.920
0.842
1.094
1.029
0.994
0.973
0.958
0.953
0.951
0.954
0.963
0.980
1.009
1.060
1.151
0.913
1.038
0.952
0.768
0.882
0.808
1.048
0.985
0.950
0.929
0.917
0.910
0.908
0.910
0.919
0.935
0.964
1.013
1.102
0.875
0.997
0.918
0,733
0.838
0,764
0.990
0.929
0.896
0.876
0.862
0.857
0.854
0.858
0.866
0.881
0.909
0.957
1.045
0.833
0.955
0.887
0.697
0.791
0.716
0.925
0.864
0.832
0.811
0.800
0.794
0.792
0.795
0.802
0.817
0.844
0.891
0.976
0.783
0.906
0.855
0.644
0.725
0.652 0.836 0.779 0.747 0.729 0,717 0.712 0312 0.714 0.721 0.737 0362 0.806 0.890 0.721 0.844 0.806
0.792
0.890 0.806
1.029
0.962
0.927
0.907
0.895
0.888
0.888 0.890
0.899
0.915
0.941
0.986
1.076
0.869
1.000
0.943
12.670 14.517 13.282 17.241 16.211 15.644 15.304 15.104 14.997 14.974 15.012 15.148 15.400 15.854 16.626 18.056 14.326 16.296 15,005
POWER PROFILE FOR EACH AVERAGE STRUCTURE
1302
1312
1322
1332
1342
Axial1
0.581
0.636
0,490 0,531
0.463
Axial2
0.513
0.571
0,440 0.482
0.425
Axial3
0.624
0.703
0.545
0,604
0.539
Axial4
0.727
0.826
0.645
0.718
0.646
Axial5
0.821
0.937
0.733
0,819
0.738
Axial6
0.901
1.028
0.805
0.900
0.812
Axial 7
0.953
1.087
0.848
0.947
0.848
Axial8
0.998
1.139
0.890
0.992
0.888
Axial9
1.019
1.162
0.906
1.009
0.900
1.026
1.169
Axial 10
0.911
1.011
0.899
Axial 11
1.022
1.165
0.908
1.008
0.897
Axial 12
1.006
1.146
0.893
0.992
0.882
Axial 13
0.984
1.122
0.878
0.979
0.876
Axial 14
0.940
1.075
0.842
0.939
0.843
Axial 15
0.886
1.017
0.799
0.897
0.810
Axial 16
0.822
0.950
0.750
0.849
0.776
Axial 17
0.740
0.863
0.687
0.784
0.725
Axial 18
0.918
1.053
0.837
0.945
0.868
Sum
15.479 17.648 13.804 15.407 13.837
POWER PROFILE FOR EACH HOT STRUCTURE
1402
1412
1422
1432
1442
0.937
1.010
0.832 .0.352
0.318
0.788
0.864
0.711
0.453 0.454
0.886
0.975
0.792
0.700
0,734
0.982
1.098
0.880
0.903
0.933
1.078
1.223
0.965
1.057
1.094
1.046
1.182
1.230
1.168
1.313
1,218
1.381
1.102
1.263
1.288
1.256
1.432
1.134
1.333
1.363
1.244
1.416
1.109
1.389
1.421
1.225
1.388
1.081
1.389
1.407
1.209
1.375
1.073
1.399
1.427
1.409
1.349
1.047
1.379
1.181
1.353
1.377
1.162
1.308
1.015
1.108
1.254
0.966
1.309
1.342
1.037
1.187
0.922
1.251
1.289
0.979
1.106
0.866
1.192
1.234
1.115
1.169
0.917
1.028
0.814
1.152
1.258
1.017
1.216
1.245
19.527 21.967 17.372 20.237 20.734
Figure 5-6: End-of-life power profiles in kW per node volume, where the node volume is constant. Top: power profile for each
plate across the core-averaged fuel element. Bottom left: power profiles for each average heat structure for input into RELAP,
normalized for one plate each. Bottom right: power profiles for each hot heat structure for input into RELAP.
302
0.221
58.16
Ave. 4th
312
0.252
11.85
Ave. 3rd
322
0.197
9.268
Ave. 2nd
332
0.220
10.34
-
Ave. End
342
0.198
9.291
Hot inner
hot 4th
402
412
0.279
0.314
0.279
0.314
1.26
1.24
Hot 3rd
422
0.248
0.248
1.26
Hot 2nd
Hot End
432
0.289
0.289
1.31
442
0.296
0.296
1.50
Total
100%
76
-
-
Ave. inner
-
Table 5.4: EOL Hot channel factor, and power generation (% of total core power)
per plate type and per lumped heat structure for RELAP input.
Plate type
Power per plate Lumped Power Hot channel factor
(1412) in the MOL core, and then to the plate 2nd from the end (1432) in the EOL
core. The hot plate in BOL case was found to be the most limiting.
The bottom axial nodes in each fuel plate typically have slightly higher power
generation due to extra neutron moderation, as can be seen in Figures 5-4, 5-5 and
5-6. It was assumed that the power generation is symmetric within each plate such
that the temperature peaks at the mid-point between the two outer surfaces. This
is a reasonable assumption regardless of the power generation in neighboring plates
because the plates are so thin.
5.2.1
Simulated Conditions
For each of the cores (BOL, MOL and EOL) both steady-state and transient conditions were simulated. Steady-state analysis is required because it shows what temperatures the coolant channel walls experience at nominal operating conditions. There
must be a 20% margin between the ONB-limiting power level and the licensed power
level, and another 20% margin between the licensed power level and the LSSS power.
Transient analysis is also important in order to observe behavior during a LOF accident and resulting temperature transient. The failure limit during an accident is the
fuel blistering temperature of 365 C, so as long as that temperature is not reached,
the clad should retain its integrity.
Steady-State
Steady-state operation at 7.0 MW with 2200 gpm (137 kg/s) flow was modeled in
RELAP5 MOD3.3 (see key parameters listed in Table 5.5) for each of the three cores
(BOL, MOL and EOL). The purpose of this simulation was to find the steady-state
temperature profiles for each unique type of channel and the hot channels in each of
the three cores studied. There must be at least a 40% margin to ONB at steady-state
operation in order for 7.0 MW to be a feasible power level for nominal operation with
the new LEU fuel.
77
LOF Transient
After steady-state operation at 7.0 MW, a RELAP-simulated LOF accident is initiated by a primary coolant low flow scram caused by simultaneous failure of both
primary pumps. In this scenario, the steady-state flow is already the low primary flow
setpoint of 2200 gpm, so the reactor scram happens at approximately the same time
as the pumps failure, as shown in Figure 5-7. As the flow rate decreases according
to the pump coastdown curve as shown in Figure 2-7, the NCVs and ASVs open,
establishing natural convection within the core tank. The fuel and clad temperatures
are of interest in the transient analysis to verify that the fuel blistering temperature
of 365 C is not reached at any point.
5.3
Beginning-of-Life
Refueling operations indicate the start of each fuel cycle. The beginning-of-life (BOL)
core is the most conservative stage to analyze because the power is concentrated in
a smaller volume (absorbers are lower in the core to compensate for added excess
reactivity) so the temperatures are higher. If the last HEU core were to be completely
replaced (contrary to normal refueling procedure) with the new 19-plate unfinned
LEU fuel, the new core is referred to as BOL, as opposed to BOC. A BOL core
represents the highest power density the fuel will ever have because the power is not
only concentrated toward to bottom half of the core height, but also in a fewer number
of elements (to compensate for the excess reactivity of all new elements). The BOL
core analyzed for 19B25 fuel performance contained only 22 fuel elements.
Steady-State Analysis
The steady-state analysis shows that even in the most limiting hot channel (442, the
hot end channel) there is at least a 22 C margin at the hottest node between wall temperature and saturation temperature (see Figure 5-8). The steady-state temperature
gradient across the fuel clad was typically approximately 1C, and the temperature
gradient from the fuel maximum to the wall temperature was typically less than 10 C.
78
Table 5.5: 19B25 steady-state conditions for all cores.
Power level (MW)
7.0
Core outlet temperature ( C)
56
Core inlet temperature( C)
43
Mass flow rate (kg/s)
137
Core pressure drop (kPa)
23.8
1.0
- - - Fraction of steady-state
-
0.8
Fraction of steady-sta
nass flow
power
S0.6
o0.4
0.2
-10
0
10
20
30
Time after SCRAM (s)
40
50
Figure 5-7: Fractional power and flow after the LOF scram
79
In the hottest average channel (312, the average 2nd channel) the margin between
wall and saturation temperature is at least 35 C. Table 6.1 shows the agreement of
various key RELAP5 MOD3.3 simulation data with the semi-analytical model (see
Chapter 6). While the hottest of the BOL hot channels is the end channel (442), the
hottest of the average channels is 312 (the average 2nd channel) and 342 (the average
end channel) is the coolest. At 7.0 MW the core temperature gradient is 13*C with
an outlet temp of approximately 56 C.
LOF Transient
Figure 5-9 shows the bulk, wall and saturation temperature in each unique type of
channel during a LOF transient. When the scram occurs, there is a temperature spike
that lasts less than 1 second while the flow rate is dropping and the reactor power
is decreasing. As natural circulation is established and reactor power continues to
decrease, there is another slower temperature rise that lasts less than 20 seconds,
before the temperature decreases monotonically with time.
The temperature rise during the LOF transient in the hot channel (442, the hot
end channel) and hot plate (1442, the hot end plate) was approximately 13 C, as
listed in Table 5.6. In the average channels, the temperature rise was less than 12 C.
The maximum fuel temperature reached during the transient is 106 C in the end
plate, and the maximum clad temperature reached is 98 C also in the end plate. The
fuel blistering temperature (and therefore failure point) is approximately 365 C, so
a maximum fuel temperature of 106 C is well within safety margins. At 7.0 MW
the margin between the maximum wall temperature and the saturation temperature
during the transient was at least 13 C in the hottest channel and at least 24*C in
the average channels. This means that for the most conservative core (BOL, only 22
elements) during a LOF event after operation at 7.0 MW the coolant will not boil
even in the hottest part of the core.
80
Ave. Inner Channel (302)
2d
Ave. 3d Channel (322)
Channel (312)
Ave. 0'
Channel
Ave.
(332)
End
Channel
(342)
,
Ave.
i
{
I
i
1
k
i
I
11
t
\
i
j
15
i
{
-
i,
I
i
i
1
i
10
i
i
i
i
I
11
I
i
i
i
i
i
i
r
I
J
7
i
J(
i
f
+l
l
0
40
60
80
100
1: 0
4
D
Temperature ('C)
60
80
100
is
0
40
Temperature ( C)
Hot Inner Channel (402)
Hot 2d
80
100
120
40
Temperature (*C)
60
80
Temperature
Hot 3'" Channel (422)
Channel (412)
20
60
Hot
4'
Channel
100
120
40
60 \
( C)
80
Temperature
100
1.
('C)
Hot End Channel (442)
(432)
I
i
i
t
I
i
i
15
0c
I
f
0
i
10
i
1
f
f
r
i
i
5
f
i
1
t
i
I
I
40
60
80
100
Temperature (*C)
120
40
60
t
80
Temperature
100
('C)
120
40
60
80
100
Temperature (*C)
120
40
60
80
Temperature
100
('C)
120
40
60
80
100
120
Temperature ('C)
Figure 5-8: BOL steady-state temperature profiles for 19B25 fuel. The solid line is the
wall temperature and the dotted line is the saturation temperature. The minimum
margin between the wall and saturation temperatures of 22 C occurred in the hot
end channel (442).
81
Average Inner Channel (302)
1 20
Hot Inner Channel (402)
U 1 00
80
--
60
----------..-
11'ff
40
Average
1 20
4
'h Channel (312)
Hot 4 " Channel (412)
E 1 00
-
-
80
1
40
Average 3rd Channel (322)
20
Hot 3"d Channel (422)
U 100
0
F
80
It
60
-------------------------------------------
40
Average 2"d Channel (332)
1 20
Hot 2 d Channel (432)
1 00
80
E
60
1
-.
40
Average End Channel (342)
H
Hot End Channel (442)
100
C.
80
60 -
I
-10
0
10
20
30
Time after SCRAM (s)
40
-10
0
10
Time after
20
SCRAM
30
(s)
40
Figure 5-9: BOL wall (solid), bulk (dashed) and saturation (dotted) temperatures for
LOF transient at 7.0 MW
82
Table 5.6: Maximum temperature rise ( C) during LOF transient at hottest node of
the wall, clad inner surface and fuel centerline in each plate type in each of the 3
cores.
Beginning-of-life
Plate type
Middle-of-life
End-of-life
Wall
Clad
Fuel
Wall
Clad
Fuel
Wall
Clad
Fuel
Ave. inner
Ave. 4th
302
10.6
10.0
9.0
9.4
8.9
8.6
14.7
14.2
10.6
312
10.4
10.3
10.0
9.9
9.9
Ave. 3rd
Ave. 2nd
322
11.5
10.9
8.5
10.6
10.0
9.6
8.0
13.5
17.0
12.9
16.5
9.3
14.1
332
11.2
10.3
9.2
10.4
9.7
8.4
14.3
13.7
11.1
Ave. End
342
10.0
9.2
8.6
9.8
10.1
7.5
12.6
12.5
8.4
Hot inner
402
11.9
11.9
11.4
11.6
11.6
11.2
11.9
11.3
10.4
hot 4th
412
11.9
12.2
11.7
11.8
12.0
11.6
11.6
11.5
11.1
Hot 3rd
Hot 2nd
422
432
14.4
13.9
13.8
12.5
10.8
10.8
14.1
13.7
11.0
16.4
15.7
13.3
12.0
11.1
10.3
12.4
11.7
11.3
Hot End
442
13.2
13.0
12.9
10.6
11.4
9.5
13.1
12.6
11.3
5.4
Middle-of-Life
The MOL core represents the BOC of an equilibrium core with partly burned fuel
but no xenon.
The power profiles for this type of core were prepared by running
depletion calculations using MCODE. The MOL core was expected to have a lower
power density due to the greater number of elements (24 as opposed to the 22 elements
in the BOL core) and therefore have slightly lower temperatures. In the MOL core
the hot channel is 412 (2nd hot channel) and on average the 2nd channel (312) is still
hotter than the other channel types.
Steady-State Analysis
Steady-state operation of the MOL core was modeled in RELAP5 MOD3.3 with
the same conditions as for the BOL case (see Table 5.5). The steady-state analysis
shows that even in the most limiting hot channel (412, the hot 4th channel) there is
at least a 26 C margin at the hottest node between wall temperature and saturation
temperature (see Figure 5-10). For the average channels, the margin between wall and
saturation temperature is approximately 36 C. The steady-state temperature gradient
83
Ave. Inner
Channel
(302)
Ave. 2 d Channel (312)
Ave.
3 d Channel
(322)
Ave.
4"
Channel
(332)
Ave.
End
Channel
(342)
t
i
i
1
I
1
20 i
1
{
t
I
I
j
{
r
I
I
1
1
I
j
j
j
1
a
I
I
3
i
I
t
15
i
1
1
I
a
i
i
I
ikk
i
II
I
E
2 10
i
I
i
i
r
r
i
5
I
i
1
r
r
r
r
i
l
1
t
J
((1
!r
,
f
ti
1
,t
t'
i
t
l
0
0
60
80
100
120
40
60
80
100
Temperature ( C)
Temperature ( C)
Hot Inner Channel (402)
Hot 2d Channel (412)
120
40
60
80
Temperature
100
120
4D
( C)
60
80
100
1
40
Temperature ( C)
Hot 3"' Channel (422)
to'
80
Temperature
Hot 46 Channel (432)
100
12
( C)
Hot End Channel (442)
20
15
0
/
1
,
10
(
1
40
60
80
100
Temperature (*C)
120
40
60
80 100
Temperature ( C)
120
40
60
80
100 120
Temperature ( C)
40
60
80
100 120
Temperature ( C)
40
60
80
100
120
Temperature ("C)
Figure 5-10: MOL steady-state temperature profiles for 19B25 fuel. The solid line is
the wall temperature and the dotted line is the saturation temperature. The minimum
margin between the wall and saturation temperatures of 26 C occurred in the hot 4th
channel (412).
84
A
Average Inner Channel (302)
120
Hot Inner Channel (402)
U 100
80
80
E
H 60
/
1
Average 4' Channel (312)
120
Hot 4 Channel (412)
100
a
s,
80
-
7,
60
AA
0.
Average
42
1 20
Channel (322)
Hot
Channel (332)
Hot 2"d Channel (432)
Average End Channel (342)
Hot End Channel (442)
3rd
3rd
Channel (422)
H
1 00
80
60
40
Average 2
1 20
G
100
80
0
H
60
40
120
U100
2
/7
I
60
40
-10
..............
7-
/'
/
r
-
0
0
10
20
30
Time after SCRAM (s)
40
-10
0
10
20
30
Time after SCRAM (s)
40
Figure 5-11: MOL wall (solid), bulk (dashed) and saturation (dotted) temperatures
for LOF transient at 7.0 MW
85
across the fuel clad was typically less than PC, and the temperature gradient from
the fuel maximum to the wall temperature was typically less than 6*C.
LOF Transient
During the temperature transient the hot channel (412, the hot 2nd channel) wall,
clad and fuel centerline temperatures experience a spike of approximately 12 C at the
hottest node, as shown in Table 5.6. The average channels experience a temperature
increase of approximately 10 C. As shown in Figure 5-11, during a LOF transient
there is margin of at least 15*C between the wall and saturation temperature at the
hottest node in the hottest channel (412, the hot 2nd channel). A MOL core with
the new LEU fuel would not experience any boiling even in the hottest point in the
core during a LOF accident.
5.5
End-of-Life
EOL represents the end-of-cycle (EOC) of an equilibrium core with deeply burned
fuel and xenon present. For this core, the hot channel was 432 (the 4th hot channel)
and on average, channels of type 312 (2nd average channel) were hotter than the
other channels by less than 2 C.
Steady-State Analysis
The steady-state EOL core was modeled with the same conditions as the BOL and
MOL cores, as shown in Table 5.5. The temperature gradient across the fuel clad
was less than 80, and the temperature gradient across the fuel was at least 1* for
all average and hot plates. Analysis of steady-state operation shows that there is a
margin of at least 24 C between the wall temperature and the saturation temperature
at the hottest node of the hottest channel (412, the 2nd hot channel), as shown in
Figure 5-12. In the average channels, a minimum margin of 31 C occurs in 312 (the
2nd average channel). With all coolant channel wall temperatures being well below
the saturation temperature, no ONB was observed.
86
Ave. Inner Channel (302)
Ave. 2"d Channel (312)
Ave. 3 rd Channel (322)
Ave.
4
*
Ave. End Channel (342)
Channel (332)
20
U
U
15
0O
00
-
10
8
Teprtr
Ho"
10 1
*)
1
hnnl(1)
5/j
0
60
80
100
12 0
Temperature (*C)
Hot Inner Channel ($02)
4
0
40
60
80
100
120
40
Temperature (*C)
Hot 3'd Channel (422)
60
80
100
120
40
Temperature (*C)
Hot 4* Channel (432)
60
80
tOO
I:
Temperature (*C)
Hot End Channel (442)
/
40
20
Q115
f
0
/
10
S
/,
40 6
0
0
2
./.
Teprtue(C
0
40
60
80
100
Temperature ('C)
120
40
60
80
100
Temperature ("C)
120
40
60
80
100
Temperature (*C)
120
40
60
80
100
120
Temperature (*C)
Figure 5-12: EOL steady-state temperature profiles for 19B25 fuel. The solid line is
the wall temperature and the dotted line is the saturation temperature. The minimum
margin between the wall and saturation temperatures of 24 C occurred in the hot 4th
channel (412).
87
Average Inner Channel (302)
120
Hot Inner Channel (402)
U 100
80
E
v/"~ ~ ----.
S60
age 4 Channel (312)
120
Hot 4" Channel (412)
U 100
80
7
CL
/
/
,-60
-.
AA
Average 3'd Channel (322)
120
Hot
3 d
Channel (422)
U100
0
E
F
80
!s //
7
60
40
Average 2"d Channel (332)
120
Hot 2"" Channel (432)
U100
80
80
E
H 60
A-
40
Average End Channel (342)
120
Hot End Channel (442)
U100
........................... . . . ... .
80
60
--
---.-
--------------------Al'
-10
0
10
20
30
Time after SCRAM (s)
40
-10
0
10
20
30
Time after SCRAM (s)
40
Figure 5-13: EOL wall (solid), bulk (dashed) and saturation (dotted) temperatures
for LOF transient at 7.0 MW
88
Loss-of-Flow Transient
During a LOF event the transient causes a temperature spike of approximately 12 C
for the hot channel (412) surface and hot plate (1412) centerline at the hottest axial
node, as shown in Table 5.6. The average channels and plates experience a temperature increase of less than 17*C. The maximum fuel temperature reached during the
transient was 97 C which occurred in the hot plate (1432, the 4th hot plate), well below the fuel blistering temperature of 365 C. The maximum clad temperature reached
was 93.7 C, which is well below the aluminum softening temperature of 450 C. The
maximum wall temperature reached was 93.5*C, as shown in Figure 5-13, leaving a
margin of at least 14 C between the wall and saturation temperatures in the hottest
channel at the hottest node.
As with the BOL and MOL cores, approximately 20
seconds after the LOF scram is initiated, all temperatures decrease monotonically
with time.
5.6
ONB-Limiting Power Level
The ONB-limiting power level was identified by the discontinuity in the heat transfer
coefficient as a function of steady-state power level. The discontinuity represents the
transition from single-phase convective heat transfer to boiling heat transfer. Because
boiling permits better heat transfer from the plate to the coolant, the heat transfer
coefficient increases suddenly at the point of ONB. Figure 5-14 shows that the 19B25
BOL ONB-limiting power level is 11.0 MW. For the MOL and EOL cores, the ONBlimiting power level is 12.5 MW.
Taking the most limiting case, the BOL core, with an ONB-limiting power level of
11.0 MW, the maximum permissible LSSS power level was calculated. As discussed
in Section 2.2, the Limiting Safety System Settings (LSSS) are based on ONB. The
margin between the LSSS and licensed steady-state power levels is currently 20%; the
LSSS power level is 6.0 MW and the licensed power level is 7.2 MW. The licensed
power level has an additional 20% margin to ONB; the ONB-limiting power level for
89
*1
24
-22
BOL
MOL
- -- EOL
I
- 18
16-
0
2
4
6
8
10
12
14
Power Level (MW)
Figure 5-14: Heat transfer coefficient at the hottest node in the hot channel for BOL,
MOL and EOL. The discontinuities indicates the ONB-limiting power level in each
of the three cases; 11.0 MW for BOL, and 12.5 MW for MOL and EOL.
the current HEU fuel is 8.4 MW. For the 19B25 LEU-fueled BOL core, the LSSS
power level would be 7.86 MW, leaving a 20% margin to the tentative licensed power
level of 9.42 MW, and another 20% margin to the ONB-limiting power level of 11.0
MW. With a maximum LSSS power of 7.86 MW, uprate to 7.0 MW nominal operating
power is feasible.
5.7
Summary of RELAP Results
The 19B25 fuel type performs well at the proposed uprate power level of 7.0 MW
with sufficient margin to ONB, even in the extreme case of the completely fresh, high
power density BOL core. At 7.0 MW the core temperature gradient is 13 C with an
outlet temp of 56 C. The maximum wall temperature reached was 83 C in the hot end
channel (442) of the BOL core during steady-state operation, with a margin of 22 C to
ONB. On average, across the three cores, the 2nd channel (312) was the hottest with
a margin of at least 35 C to ONB. The maximum clad temperature at steady-state
was 85 C and the maximum fuel temperature was 93 C in the hot end plate (1442)
of the BOL core. The fuel centerline-to-wall temperature gradient was less than 10 C.
90
During the first few minutes of a LOF transient, the maximum fuel temperature
reached was 106 C in the hot end plate (1442) after a rise of approximately 13 C
from steady-state conditions at 7.0 MW. During the transient the minimum margin
to ONB was at least 13 C in the hot channel and approximately 24 C in the average
channels. A summary of the hot channel maximum temperatures for each core is provided in Table 5.7. These temperatures are all well below the aluminum clad softening
temperature of 450 C and fuel blistering temperature of 365 C, so the 19B25 fuel type
appears to perform well during LOF events after nominal operation at up to 7.0 MW.
Table 5.7: Hot channel maximum temperatures for each core.
Core
Hot Plate
Transient
Steady-state
Wall
Clad
Fuel Wall
Clad
Fuel
BOL
1442
83.2
84.6
93.2
96.4
97.6
106.0
MOL
1412
82.0
82.5
88.1
93.8
94.5
99.7
EOL
1432
81.1
82.0
85.4
93.5
93.7
96.7
91
92
Chapter 6
2D Semi-Analytical Validation
A semi-analytical steady-state 2D lumped heat structure model was created in Mathematica to verify the RELAP5 MOD3.3 results. Calculated parameters such as the
hydraulic diameter, Reynolds number and heat transfer coefficient were used to validate the RELAP input. Maximum steady-state surface, clad and fuel temperature
results calculated from the model were compared with RELAP results and found to
agree within a relative error of 8%.
The hydraulic diameter De was calculated as shown in Equation 6.1, where Af is
the channel flow area, and PW is the wetted perimeter.
De =
4Af
(6.1)
PW
The flow area of each lumped channel was calculated as shown in Equation 6.2,
where wg is the gap width, wP is the plate width, and nch is the number of physical
channels. For hot channels, wP is instead the width of the fuel wf, as explained in
Section 2.2.2.
Af = wg x wp x nch
(6.2)
The wetted perimeter of each lumped channel was calculated using Equation 6.3.
For hot channels, wP is fuel width wf.
93
Pw = rch x 2 (wg
+ wp)
(6.3)
For high aspect ratio geometries, the hydraulic diameter can be reasonably approx-
imated as twice the wetted perimeter, though Equation 6.2 was used for all hydraulic
diameter calculations.
The mass flow rate through each channel was obtained by solving the mass continuity and momentum equations. The flow split among coolant channels was determined by the pressure difference between the inlet and outlet plenum of the core as
shown in Equation 6.4, which is dominated by the friction pressure drop in Equation
6.5. An analytical solution was derived for the mass flow rate ratio between nominal
(average) coolant channels and an abnormal (hot) coolant channel for the MITR, as
shown in Equation 6.6. [5]
AP = f
(6.4)
De 2
4
f= 0.316
)
pv
12
(i
kIl2
=
(wp,2wg,27
(
wp,iwg,i
(6.5)
5
wp,lwg,1
wp,2wg,2)
(6.6)
For average channels where the plate width is the same but the gap width is
different, Equation 6.6 is reduced to Equation 6.7, which assumes that wg « wp
which is true in the case of MITR fuel. [5]
12
\m12/
wg,2/
(6.7)
The Reynolds number was calculated from Equation 6.8, where pw is the viscosity
of water at the inlet temperature (approximately 6 x 10-4 Pa s at 43 C).
94
Re =
rhi Dei
(6.8)
The Dittus-Boelter correlation was used to find the Nusselt number for convective
heat transfer at the fuel plate surface, as shown in Equation 6.9. Re is the Reynolds
number as calculated using Equation 6.8 and Pr is the Prandtl number as calculated
using Equation 6.10, where cp and k, are the specific heat and thermal conductivity
of water respectively.
Nu = 0.023 Re0. 8 Pr0 .4
Pr = c p
(6.9)
(6.10)
Using the Nusselt number calculated in Equation 6.9 and the hydraulic diameter
from Equation 6.1, the heat transfer coefficient at the plate surface was calculated,
as shown in Equation 6.11.
h
Nu k,
D=
De
(6.11)
The model is semi-analytic because the integration in the direction of heat transfer (from the fuel meat to the plate surface) is analytical, while the integration in the
direction of coolant flow (vertical, parallel to the plate surface) is numerical because
the power generation profile extracted from MCODE (see Section 2.3) is discrete with
18 nodes.
Equations to model radial heat transfer and axial temperature were derived starting with Equation 6.12 where k is the thermal conductivity and q "' is the volumetric
heat generation rate. The term for temperature change with time was eliminated
because only steady-state conditions were to be studied. The gradient term was expanded and simplified to contain a positional derivative in only one direction (from
fuel to plate surface).
95
p cpdT = kV 2T + q"'
(6.12)
dt
62
0=k
2
2
82+
+
z2
T
q
d kdT 1
dx dx
(6.13)
0 = k--- + q"'x
dx
With a reflective boundary condition at the fuel centerline, Equation 6.13 was
solved for each layer from the fuel meat to the plate surface. The thermal resistances
and thus equations that describe the temperature profile in each layer were obtained
as shown in Equation 6.14, where the variables and subscripts are as follows:
f
fuel
z
zirconium foil
c
clad
6
thickness
k
thermal conductivity (W/mK)
heat transfer coefficient (W/m2 K)
q"
heat flux (W/m 2
)
h,
TCo
clad outer (plate surface) temperature (K)
TCL
fuel centerline temperature (K)
TCL =
2kp
kz
To+q + +_(6.14)
kc
e
To model the axial temperature profiles, Equation 6.16 was derived
96
from
Equation
6.15, where cPW is the specific heat of water in units of J/kgK, mi is the mass flow rate
in kg/s, and subscript i indicates the quantity for node i of 18 axial nodes. The linear
heat generation rate in Equation 6.15 was replaced with the total heat generation
for each node, as shown in Equation 6.16 because the calculation steps through each
node, from 1 to 18.
dT.
mm cpW = q'(z)
dz
T(z) =
Ti,
+ q'z
m cpW
(6.15)
Ti=Ti-1 +AT
Ti = Ti- + .
M cPW
(6.16)
Using this model for 2D steady-state temperature distribution along a channel,
19B25 BOL results from RELAP5 MOD3.3 were verified. Figure 6-1 shows the temperature profiles for each type of channel produced with the semi-analytical model.
Table 6.1 lists mass flows, heat transfer coefficients and temperatures for the mathematical model compared to RELAP, and their relative errors.
All relative errors
were less than 7%. The source of the error is likely the coupling in RELAP that
wasn't accounted for in the semi-analytical model. For example, in the RELAP input
all the hot channels are adjacent to each other so the temperature in one channel
may influence the temperature in the next, while all the channels are decoupled in
the semi-analytical model. The semi-analytical model verifies that the geometric input for RELAP was correct and that RELAP treats the 1-phase flow heat transfer
correctly for steady-state conditions.
97
100
100
1302
1402
90
90
T
= 87.9
C-82.
7
6.
80
T-
80
-
T s= 81.6
"'*.
70
0
7.8
60
I- 00
Tb
40'
0
_
5
10
--
40
15
100
0
1312
Tf =76.8
Tf
70
S60
60
-_
0
10
5
=
84.6
50
40
____
15
..... .. . ...... . .
.. .
.......
Ts= 78.8
Tb = 61.4
Tb=56.
0
5
10
1322
1422
90
T T=
80
Tf
=
15
100
90,
i
15
80
70
10 0
,69.8
10
1412
50
.
5
90
80
40
62.7
100
90
v
=
50
80.3
73.6
701
70'
76.3
-
=
E,
60
60
50
40.
0
5
10
40
15
100,
90
0
15
10
5
1001332
1432
90
801
Tf = 84.5
80-
Tf= 73.9
S70'
70
=
T=7.
0.
v 60
t50'
60
Tb
40'
0
10
5
=
Tb = 60.3
50
50
55.1
40-
15
0
100
10
5
15
100
1342
901
a
Tb =60.1
50,
Tb =55.1
90'
l442
-----
87.6
80
80
T = 70.8
70
9
T= 67.5
-
wati
60
50!
40
0
TOWa
82.
70'
60T
TI
Tmax
c-
50,
Tb
=
Tb = 53.2
50.2
40
0
5
10
15
Axial Node (1 = bottom, 18 = top)
5
Axial Node (1
=
10
bottom, 18
=
15
top)
Figure 6-1: Bulk, wall, clad and fuel temperature profiles in each type of channel.
98
Table 6.1: Comparison of RELAP results with semi-analytical model for 19B25 BOL.
Quantity
Plate Type
Mass flow (kg/s)
Ave. inner
Ave. 4th-2nd
Ave. end
Hot inner
Hot 4th-2nd
Hot end
Ave. inner-2nd
Ave. end
Hot inner-2nd
Hot end
Ave. inner
Ave. 4th
Ave. 3rd
Ave. 2nd
Ave. end
Hot inner
Hot 4th
Heat transfer
coeff. (kW/m2 K)
Peak wall temp. ( C)
Peak clad temp. (0C)
Structure
Model
RELAP
Rel. Error (%)
302
312-332
342
402
412-432
442
302-332
342
402-432
442
302
312
322
332
342
402
412
81.7
13.4
10.8
0.28
0.28
0.23
15.6
16.0
15.6
16.0
70.8
71.5
69.8
70.0
67.5
81.6
78.8
78.7
12.9
10.4
0.27
0.27
0.22
15.9
15.8
16.3
16.3
71.4
74.7
69.6
71.8
69.1
80.5
81.7
-3.7
-3.6
-3.4
-2.7
-3.9
-2.6
+1.9
-1.3
+4.5
+1.9
+0.8
+4.5
-0.2
+2.6
+2.4
-1.3
+3.7
Hot 3rd
422
76.3
75.6
-0.9
Hot 2nd
Hot end
432
442
302
312
322
79.4
82.2
71.4
72.4
77.8
83.2
72.0
75.5
-2.0
+1.3
70.5
332
70.8
342
402
412
422
432
442
302
312
322
68.4
79.5
77.1
80.4
83.6
76.0
76.8
73.6
73.9
70.8
87.9
84.6
80.3
70.3
72.8
69.9
81.3
82.3
76.2
79.3
84.5
76.7
80.8
73.4
76.1
75.5
86.8
88.2
79.5
84.5
82.6
87.6
93.1
Ave. inner
Ave. 4th
Ave. 3rd
Ave. 2nd
Ave. end
Hot inner
Hot 4th
Hot 3rd
Peak fuel temp. ( C)
Hot 2nd
Hot end
Ave. inner
Ave. 4th
Ave. 3rd
Ave. 2nd
332
342
402
412
422
432
442
Ave. end
Hot inner
Hot 4th
Hot 3rd
Hot 2nd
Hot end
99
82.4
+0.8
+4.3
-0.3
+2.8
+2.2
-1.3
+3.5
-1.2
-1.4
+1.1
+0.9
+5.2
-0.3
+3.1
+6.6
-1.3
+4.3
-1.0
-2.2
+6.3
100
Chapter 7
Burnup Effect on Fuel
Temperature
There are two potential mechanisms that affect the peak fuel temperatures analyzed
in Chapters 4 and 5. The first is oxide layer buildup on the cladding surface and the
second is the changes in the U-1OMo fuel meat. Both can be correlated to fuel burnup
(fission density). Because of the complex refueling schemes and core configurations
of the MITR-II, bounding analyses were performed to evaluate the effect of oxide
layer and fuel thermal conductivity on the peak fuel temperatures. All analyses were
performed for BOL, which was shown in Chapter 5 to be the most limiting in a fuel
cycle.
As fuel is irradiated in the core, oxide forms on the cladding interface with the
coolant and the material composition of the fuel changes with burnup. The oxide layer
has a lower thermal conductivity than the clad and therefore reduces heat conduction,
and the thermal conductivity of the fuel is reduced with neutron fluence. The effect
of oxide formation and fuel burnup on fuel and clad temperatures are quantified in
the following sections.
101
7.1
Effect of Oxide Layer
Over the in-core life of a fuel element an oxide layer forms on all cladding surfaces
as a product of corrosion. The formation of the oxide layer is primarily a result of
corrosion of the aluminum cladding. The reaction rate of the oxidation process is
a function of temperature, heat flux, flow velocity, and chemical (pH) conditions.
Previous analyses conservatively recommends that an oxide layer thickness of 2 mils
(0.0508 mm) be adopted based on the limit to prevent spallation that could lead to
fission product gas release. [24] This oxide layer has a lower thermal conductivity than
the aluminum clad and therefore reduces the effective thermal conductivity, raising
the peak fuel temperature. The purpose of this semi-analytical sensitivity study was
to quantify the effect of oxide on the wall temperatures during steady-state operation
of a 19B25 LEU core.
Using the semi-analytical model discussed in Chapter 6, an oxide layer was added
to each fuel plate. Equation 7.1 shows the expression for the fuel centerline temperature including the oxide thermal resistance term with subscripts "o". The oxide layer
was 0.002" (0.0508 mm) thick and had a thermal conductivity of 2.25 W/mK. [11] It
was assumed that the oxide layer was so thin that it did not displace clad nor coolant
channel flow area.
TCL=Teo
q
if
C
Z
2krkz
c
e
6
1
(7.1)
o he
With the addition of a 0.0508 mm oxide layer to the surface of each fuel plate
in the BOL case (using BOL power profiles at 7.0 MW), the steady-state clad and
fuel temperatures experienced an average increase of approximately 90, as shown in
Table 7.1 and Figure 7-1. The maximum clad and fuel temperatures reached were
96 C and 100 C respectively. These temperatures are well below the fuel blistering
102
temperature of 365*C and the clad softening temperature of 4500, and therefore any
oxide build-up on the plate surfaces over the life of the core would not cause the clad
and fuel temperatures to exceed the temperature limits.
Table 7.1: Peak fuel and clad temperatures before and after the addition of a 0.0508
mm oxide layer.
Peak clad temp. ( C)
Ave. inner
Ave. 4th
Ave. 3rd
Ave. 2nd
Ave. end
Hot inner
4th
Hot 3rd
With Oxide
302
312
322
71.4
79.3
72.4
70.5
80.2
78.1
332
342
70.8
78.3
75.8
91.9
402
412
68.4
82.4
422
79.5
77.1
Hot 2nd
432
80.4
Hot end
442
Ave . inner
302
312
Hot
Peak fuel temp. (*C)
Without Oxide
Ave . 2nd
332
Ave . end
342
Hot inner
402
4th
Hot 3rd
412
422
83.6
76.0
76.8
73.6
73.9
70.8
87.9
84.6
80.3
Hot 2nd
432
84.5
Hot end
442
87.6
Ave .
4th
Ave . 3rd
Hot
322
103
AT
7.9
7.8
7.6
7.5
7.4
9.5
88.3
8.8
85.0
90.3
95.6
83.9
84.8
7.9
9.9
81.2
81.6
78.2
97.5
93.5
88.4
94.5
99.5
12.0
7.9
8.0
7.6
7.7
7.4
9.6
8.9
8.1
10.0
11.9
120
120
1302
100
Tf
5
EHe
T j = 97.5
= 83.9
80
80
=
60
70.8
2816
-
U
1402
100:
601
Tb = 62.7
Tb = 55.9
40
40
0
5
10
15
0
120
1312
U
100
15
Tj93.5
= 84.8
80
80
T, = 78.8
T=71.5
60
60
=61.4
Tb = 56.
40
40
0
5
10
15
0
120
5
10
15
120
1322
C
10
1412
100
Tj
0
5
120
1422
100
100
80
80
60
60
Tf = 88.4
E
E=
T,=76.3
Tb = 60.1
40
Tb =55.1
0
5
t
20
10
15
0
1332
U
5
10
15
120
-
-
401
1432
1 00
100
80
T
Tf = 94.5
80
=783..
;a
a.
H~
60
60
Tb = 60.3
--]015
40
T6=55.1
0
5
10
40
15
120
1
120
1342
1442
U 100
82.
- Wa3.
-- 5
0
T=
100
Tf=
80
T
,
78.2
80
-
Tbulk
-
Tlad
99.5
.
=
T
75
H 60
60
40,
0
T,
5
10
Axial Node (1 = bottom,
=
50.2
Th =53.2;
40
15
0
18 = top)
5
Axial Node (1
10
=
15
bottom, 18 = top)
Figure 7-1: Semi-analytical model steady-state 7.0 MW temperature profiles for each
type of plate with a 0.0508 mm oxide layer.
104
7.2
Effect of Fuel Thermal Conductivity
Recent work on the thermal properties of irradiated mini-plate coupons has shown a
significant reduction in U-I0Mo fuel meat thermal conductivity with burnup. The empirical correlation of U-1OMo thermal conductivity as a function of temperature and
burnup is provided in Equation 7.2 where
#
is the fission density in 1021 fissions/cm 3
and T is the temperature in 0C. [11]
k = 12.57+0.04T - 0 x (1.322 + 0.00278T) - T2 (2.351 x 10- 5 +4.996 x 10- 6 0) (7.2)
Using the empirical correlation, the fresh fuel thermal conductivity is 14.28 W/mK
.
and that for depleted fuel is 7.01 W/mK with a fission density of 5 x 1021 fissions/cm3
Taking into account an estimated measurement uncertainty of approximately 20%,
the fuel meat thermal conductivity is reduced to 5.61 W/mK at worst.
[4] This
conservative value for the irradiated fuel thermal conductivity was adopted for the
bounding analysis of peak fuel temperature.
The current HEU fuel has a burnup limit of 1.8 x 1021 fissions/cm 3 , and 5 x 1021
fissions/cm 3 is the proposed burnup limit for the LEU fuel. A burnup this high is the
bounding burnup case and therefore equates to the worst thermal conductivity for an
MITR LEU core.
The analytical solution for plate type fuel (assuming constant thermal conductiv-
ity in the fuel meat) is as shown in Equation 7.3, where Tm.
is the fuel centerline
temperature in Kelvin, Tci is the clad inner temperature in Kelvin, a is the fuel meat
half-thickness in meters, q" is the heat flux in (W/m 2 ), and kf is the fuel meat thermal conductivity in W/mK. [27] The temperature differential from the fuel centerline
to the clad inner surface is inversely proportional to the fuel thermal conductivity,
as can be see in Table 7.2. Created from the empirical thermal conductivity formula
in Equation 7.2. Figure 7-2 shows the thermal conductivity of the U-1OMo fuel as a
105
function of temperature for various levels of burnup. For the MOL and EOL cores, the
peak fuel temperatures calculated analytically for a fresh fuel nominal thermal con)
ductivity of 14.28 W/mK, for depleted (to the burnup limit of 5 x 1021 fissions/cm 3
thermal conductivity of 7.01 W/mK, and a worst-case (to account for 20% uncertainty) thermal conductivity of 5.61 w/mK are given in Table 7.3.
(Tmax
- Tai)
-
(7.3)
q"
The highest theoretical fuel temperature caused by reduced thermal conductivity
due to maximum burnup is 90 C, which is well below the fuel blistering temperature
of 365 C.
201
Ox1021
Depletion
,
2x10=1
3x10'
U
.... .~
~.
4x102 1
0
5x10
50
100
150
21
200
Temperature ( C)
Figure 7-2: Thermal conductivity of U-1OMo fuel meat as a function of temperature
and burnup (0p in fissions/cm 3). [4, 11]
106
Table 7.2: Burnup effect on thermal conductivity of U-10Mo fuel meat. The halfthicknesses for the 100%, 70% and 55% nominal fuel meat thicknesses are
0.32 mm,
0.22 mm and 0.17 mm respectively.
Thermal Conductivity (W/mK)
Fuel half-thickness
ALT
100%
70%
55%
14.28
2.8
1.9
1.4
7.01
5.6
3.8
2.9
5.61
7.0
4.8
3.7
Table 7.3: Burnup effect on fuel centerline temperature of U-lOMo fuel meat. For
each plate type for both MOL and EOL cores, the peak fuel temperature is provided
for unirradiated fuel, fuel irradiated to the burnup limit, and fuel irradiated to the
burnup limit with a 20% reduction in thermal conductivity.
MOL
EOL
Peak Fuel Temperature ( C)
14.28 W/mK 7.01 W/mK 5.61 W/mk
Plate Type
Tci ( C)
302
71.4
74.2
77.0
78.4
312
74.4
77.2
80.0
81.4
322
332
342
69.1
70.9
67.3
71.0
72.8
68.7
72.9
74.7
70.2
73.9
75.7
71.0
402
81.5
84.3
87.1
88.5
412
82.5
85.3
88.1
89.5
422
76.2
78.1
80.0
81.0
432
442
302
312
322
332
342
402
412
422
78.3
74.5
71.5
74.5
69.7
71.5
67.5
78.5
80.8
75.5
79.7
75.9
74.3
77.3
71.6
72.9
68.9
81.3
83.6
77.4
81.2
77.4
77.1
80.1
73.5
74.4
70.4
84.1
86.4
79.3
82.0
78.2
78.5
81.5
74.5
75.2
71.2
85.5
87.8
80.3
432
82.0
83.4
84.9
85.7
442
80.7
82.1
83.6
84.4
107
108
Chapter 8
Summary and Conclusions
HEU fuel for use in research and test reactors is being phased out globally in ac-
cordance with the goals of the GTRI program. When the Megatons to Megawatts
program ended in 2013, approximately 20,000 HEU Russian warheads had been dismantled and used to fuel research and test reactors world-wide. The supply of avail-
able HEU is diminishing and to improve global special nuclear material security, all
research and test reactors must convert to LEU.
The MITR-II currently uses HEU fuel and its neutronic performance would be
stifled, due to parasitic neutron absorption by
238
U, without conversion to a high-
density (17 g/cm 3 , non-dispersion) metallic U-10Mo LEU fuel, and without a power
uprate. In order to reduce end-plate power peaking in the fuel to make power uprate
thermal-hydraulically feasible, a fuel redesign was found to be necessary.
The first
study was to quantify the effect of changing finned fuel to unfinned. Removal of the
fins was desirable because not only does it simplify the manufacturing process, it
also permits thinner fuel clad, therefore making more room for more fuel plates per
element, thus reducing the power density per plate.
The reference case was an 18-plate LEU fuel with fins which was compared to an
identical case without fins using RELAP5 MOD3.3. The ONB-limiting power level of
the unfinned fuel reduced to approximately half that of the finned fuel, and the wall
109
temperature exceeded the saturation temperature in the hot channel in both cases at
the proposed uprate power level of 7.0 MW. While the fins did assist in cooling, the
effect of removing the fins can be compensated for by reducing the power density per
plate (increasing the number of plates also increases the heat transfer surface area)
and increasing the coolant flow rate.
Work done at ANL involved neutronic analysis of a design space comprising unfinned fuel with a greater number of plates per element, LEU fuel meat and graded fuel
meat thickness. Gradation of the fuel meat thickness allowed for more even power
density from plate-to-plate within the fuel element, thus raising the ONB-limiting
power level for the core.
Of the fuels studied, a 19-plate fuel with 25 mil nominal fuel meat thickness was
found to be a promising candidate for the conversion with an increased flow rate to
2200 gpm. The scope of this study was to characterize the thermal-hydraulic performance of this particular fuel design for various burnup levels at nominal steady-state
conditions and during a LOF temperature transient. The steady-state analysis was
necessary to find the ONB-limiting power level and to identify the wall-to-saturation
temperature margins. The transient analysis was necessary to quantify the maximum
fuel and clad temperatures reached to be certain that the softening temperature of
the clad (450*C) and the blistering temperature of the fuel (365 C) were not exceeded
at any point.
The main conclusions of this study are as follows:
* Upon analysis of BOL, MOL and EOL cores, (where the BOL core had only 22
elements while the MOL and EOL cores had 24) the BOL core was found to be
the most limiting. This was expected because with the reduced number of fuel
elements in the core, the power density per plate and therefore the temperature
gradients across the fuel were higher.
110
The ONB-limiting power level for the
BOL core was 11.0 MW, while that for the MOL and EOL cores was 12.5 MW.
Even in the most limiting case (BOL) there is more than a 40% margin between
the ONB-limiting power level and the proposed uprate power level for nominal
operation of 7.0 MW.
" In the BOL core, the hot end channel (442) was the most limiting, and the 2nd
channel (312) was the hottest of the average channels. With burnup, the most
limiting channel became the hot 2nd channel (412) in the MOL core, and on
average channels of type 312 were still hotter than the others. In the EOL core,
the 4th hot channel (432) was found to be the most limiting and channels of type
312 continued to be hotter on average than other channels. While the hot channel relocated with burnup, channels of type 312 were consistently the hottest in
each element (by 3-8 0C), meaning that the fuel meat gradation was successful
in reducing end plate power peaking. A semi-analytical 2D heat transfer model
was used to verify the BOL results, and the RELAP BOL temperatures were
found to agree within a relative error of 8%.
" In the most limiting channel in the most limiting core, the highest wall temperature was 83.2 C, with a 22 C margin to the saturation temperature.
The
maximum wall temperature in the average channels was approximately 70 C
with a 35 C margin to the saturation temperature. These steady-state results
show that no boiling occurs even in the hottest part of the hottest channel with
the highest power density per plate the LEU core would ever see, with still
enough temperature margin before ONB to permit an uprate to 7.0 MW.
" During a LOF transient initiated due to simultaneous primary pump failure,
the fuel experiences a temperature transient as a result of the loss of forced
convection and transition to natural convection cooling.
There is an initial
temperature spike as the coolant flow rate drops faster than the power does, and
then a temperature dip and slower rise again as natural circulation is established.
The temperatures drop below the steady-state temperatures within 30 seconds
of the scram. The peak fuel temperature reached during the scram was 106 C in
111
the hot end plate (1442) of the BOL core, an increase of 13 C from steady-state
conditions. This is well below the fuel failure limit (blistering temperature) of
365 C meaning that the fuel meat performs satisfactorily during a LOF accident
after steady-state operation at 7.0 MW. The channel wall temperatures during
the transient remained below the saturation temperature, even in the BOL core,
with a margin of 13 C in the hot channel (442) and approximately 35 C in the
average channels, so no boiling occurred even in the most limiting channel of
the most limiting core upon LOF.
* To further stress-test the LEU fuel, an analytical study was done to quantify
the effect of clad surface oxide buildup on clad and fuel temperatures. With an
oxide layer of 2 mil (0.0508 mm), the maximum temperature rise was 12 C to
reach a maximum fuel centerline temperature of 100 C, still well below the fuel
blistering temperature.
* The effect of burnup and temperature on fuel meat thermal conductivity was
also quantified. With fuel irradiated to the proposed burnup limit of 5 x 1021
fissions/cm 3 and a further 20% reduction in thermal conductivity to account
for the thermal conductivity empirical measurement uncertainty, the maximum
fuel temperature was 90 C, again well below the fuel blistering temperature of
365 C.
Overall the 19-plate graded fuel with 25 mil nominal fuel meat thickness is a
promising candidate for the MITR conversion to LEU fuel. The fuel provides sufficient
margin to ONB during steady-state operation to safely permit a power uprate to 7.0
MW. During a LOF, not only is there a significant temperature margin to the clad
softening and fuel blistering temperatures in the hottest plate, the wall temperature
does not exceed the saturation temperature at any point in the hottest channel,
meaning that no boiling is expected to occur. The conclusions of this study support
the new 19B25 fuel design work for the LEU conversion of the MITR.
112
Bibliography
[1] RELAP5/MOD3.3 Code Manual, March 2006.
[2] A. Bergeron, E.H. Wilson, G Yesilyurt, F.E. Dunn, J.G. Stevens, L. Hu, and T.H.
Newton. Low enrichment uranium core design for the massachusetts institute
of technology reactor (MIT) with un-finned cladding 12 mil-thick clad UMo
monolithic fuel. Technical Report ANL/GTRI/TM-13/15, Argonne National
Laboratory, November 2013.
[3] A. E. Bergles and W. M. Rohsenow.
The determination of forced-convection
surface-boiling heat transfer. 86:365-372, 1964.
[4] D.E. Burkes, A.M. Casella, E.C. Buck, A.J. Casella, M.K. Edwards, P.J. MacFarlan, K.N. Pool, B.D. Slonecker, F.N. Smith, F.H. Steen, and R.E. Thornhill. Fuel
thermo-physical characterization project: Fiscal year 2013 final report. Technical
report, Pacific Northwest National Laboratory, 2013.
[5] K. Chiang. Thermal hydraulic limits analysis for the mit research reactor low
enrichment uranium core conversion using statistical propagation of parametric
uncertainties, May 2012.
[6] D.R. Don. Natural circulation in the MITR-II, May 2014.
[7] D.R. Don. Simplified MITR-II plant, May 2014.
[8] F.E. Dunn. MNSR transient analysis and thermal hydraulic safety margins for
HEU and LEU cores using the RELAP5-3D code. In Int. Conf. on Reduced
Enrichment for Research and Test Reactors, Prague, Czech Republic, September
2007.
[9] F.E. Dunn, A. Olson, E.H. Wilson, K. Sun, T.H. Newton, and L. Hu. Preliminary accident analyses for conversion of the massachusetts institute of technology
reactor (MITR) from highly enriched to low enriched uranium. Technical Report
ANL/GTRI/TM-13/5, Argonne National Laboratory and the Massachusetts In-
stitute of Technology, July 2013.
[10] E. Feldman. Thermal properties to use in MURR accident analysis.
113
[11] E.E. Feldman, L.P. Foyto, K. Kutikkad, J.C. McKibben, N.J. Peters, J.G.
Stevens, J.A. Stillman, and E.H. Wilson. Accident analyses for conversion of
the university of missouri research reactor (murr) from highly-enriched to lowenriched uranium. Technical Report ANL/GTRI/TM-14/5, Argonne National
Laboratory and the Massachusetts Institute of Technology, 2014.
[12] E. Forrest and L. Hu.
Experimental investigation of single-phase heat trans-
fer and onset of nucleate boiling in a prototypic materials test reactor coolant
channel to support the MITR LEU conversion. Technical Report MIT-NRL-1401, Nuclear Reactor Laboratory, Massachusetts Institute of Technology, January
2014.
[13] S.J. Kim, L. Hu, and F.E. Dunn. Thermal-hydraulic analysis for heu and leu
transitional core conversion. Nuclear Technology, 182(3):315-334, June 2013.
WOS:000319638300006.
[14] Y. Ko. Themal Hydraulic Analysis of the MIT Research Reactor in Support of a
Low Enrichment Uranium (LEU) Core Conversion. PhD thesis, Massachusetts
Institute of Technology, Boston, USA, 2008.
[15] MIT Nuclear Reactor Laboratory. Reactor Systems Manual, 2004.
[16] MIT Nuclear Reactor Laboratory. MITR Users' Guide, 2012.
[17] T.H. Newton. Memorandum re: Pump coastdown, May 2011.
[18] T.H. Newton, L. Hu, G.E. Kohse, E.E. Pilat, B. Forget, P. Romano, Y. Ko,
S. Wong, Y. Wang, B. Dionne, J. Thomas, and A. Olson. Completion of feasibility studies on using LEU fuel in the MIT reactor. In RERTR, Beijing, China,
2009.
[19] T.H. Newton, M.S. Kazimi, and E.E. Pilat. Development of a low enrichment
uranium core for the MIT reactor. Nuclear Science and Engineering, 157:264-
279, 2007.
[20] NRC. 10 CFR part 50. Technical report, 1986.
[21] H. Ozaltun, R.M. Allen, and Y.S. Han. Effects of the zirconium liner thickness
on the stress-strain characteristics of u-10Mo alloy based monolithic mini-plates.
Proceedings of the AMSE 2013 International Mechanical Engineering Congress
& Exposition, IMECE2013(66595), November 2013.
[22] A.B. Robinson, G.S. Chang, D.D. Keiser, D.M. Wachs, and D.L. Porter. Irradiation performance of u-mo alloy based 'Monolithic' plate-type fuel - design selec-
tion. Technical Report INL/EXT-09-16807, Idaho National Laboratory, Idaho
Falls, Idaho, August 2009.
114
[23] K. Shirvan. The Design of A Compact Integral Medium Size PWR: The CIRIS.
Master of science in nuclear science and engineering, Massachusetts Institute of
Technology, June 2010.
[24] MITR staff. Safety analysis report for the MIT research reactor (MITR-II).
Technical Report SAR-2011, MITNRL, 2011. p. 137-138.
[25] Y. Sudo, K. Miyata, H. Ikawa, and M. Kaminaga. Experimental study of incipient
nucleate boiling in narrow vertical rectangular channel simulating subchannel of
upgraded JRR-3. 23(1):73-82, January 1986.
[26] K. Sun, M. Ames, T.H. Newton, and L. Hu. Neutronic analysis and fuel cycle
simulation of the MIT reactor using MCODE-FM and experimental validation.
Proceedings of ICONE21, (Paper 16670), 2013.
[27] N.E. Todreas and M.S. Kazimi. Nuclear Systems: Vol 1. CRC Press, 2 edition,
2012.
[28] USEC. Megatons to megawatts, 2014.
[29] D. Wachs, D. Keiser, Y.S. Kim, P. Medvedev, D. Perez, G.L. Hofman, M. Meyer,
J. Jue, M. Okuniewski, J. Gan, A. Robinson, B. Rabin, A. Wertsching, F. Rice,
C. Papesch, M. Lillo, G. Chang, H. Ozultun, N. Woolstenhulme, I. Glagolenko,
S. Miller, and P. Murray. Draft report on information relevant to u-mo fuel
design. Technical Report INL/LTD-12-25703, Idaho National Laboratory & Argonne National Laboratory, March 2013.
[30] Dan Wachs. RERTR fuel development and qualification plan. Technical Report
INL/EXT-05-01017, Idaho National Laboratory, Idaho, USA, 2007.
[31] R.H. Whittle and R. Forgan. A correlation for the minima in the pressure drop
versus flow-rate curves for sub-cooled water flowing in narrow heated channels.
Nuclear Engineering and Design, 6(1):89-99, August 1967.
[32] W.L. Woodruff, N.A. Hanan, and J.E. Matos. A comparison of the RELAP5/MOD3 and PARET/ANL codes with the experimental transient data
from the SPERT-IV d-12/25 series. 1997.
[33] Z. Xu. Design Strategiesfor Optimizing High Burnup Fuel in Pressurized Water
Reactors. PhD thesis, Massachusetts Institute of Technology, Boston, USA, 2003.
115