Development of Integral Actuation Technology for Composite
Rotorcraft Structures
by
Mads C. Schmidt
B.S., Massachusetts Institute of Technology (1998)
Submitted to the Department of Mechanical Engineering
in partial fulfillment of the requirements for the degree of
Master of Science in Mechanical Engineering
at the
MASSACHUSETTS INSTITUTE OF TECHNOLOGY
BARKER
February 2001
MASSACHUSETTS INSTITUTE
OF TECHNOLOGY
@ Massachusetts Institute of Technology 2001
All rights reserved
LIBRARIES
Signature of Author.............
Certified by .............
.........
Department of Mechanical Engineering
27 November 2000
....................
Nesbitt Hagood IV
Associate Professor, Department of Aeronautics and Astronautics
Thesis Supervisor
C ertified by ................
....................
(
JSeth
Lloyd
Assistant Professor, Department of Mechanical Engineering
Mechanical EnginesagAt%@Qty Reader
A ccepted by .........................................
Ain Sonin
Chairman, Department Committee on Graduate Students
MiTLibraries
Document Services
Room 14-0551
77 Massachusetts Avenue
Cambridge, MA 02139
Ph: 617.253.2800
Email: docs@mit.edu
http://libraries.mit.edu/docs
DISCLAIMER OF QUALITY
Due to the condition of the original material, there are unavoidable
flaws in this reproduction. We have made every effort possible to
provide you with the best copy available. If you are dissatisfied with
this product and find it unusable, please contact Document Services as
soon as possible.
Thank you.
Due to the quality of the original material there is some bleed
through.
Development of Integral Actuation Technology for Composite Rotorcraft
Structures
by
Mads C. Schmidt
Submitted to the Department of Mechanical Engineering
on 27 November 2000, in partial fulfillment of the
requirements for the degree of
Master of Science in Mechanical Engineering
Abstract
Helicopters experience a large level of vibration in flight due to the unsteady aerodynamics
associated with the spinning rotor. Reduction of vibration can improve ride comfort, reduce
cabin noise, and improve maintainability and functional reliability. Blade mounted actuation
technology using smart materials has made effective rotor vibration reduction possible.
The goal of the work presented in this thesis is to further develop integral actuation technology for rotor blades. Active Fiber Composite (AFC) plies, consisting of piezoelectric fibers
embedded in an epoxy matrix, are embedded in the skin of model scale rotor blades to allow
the blades to be actively twisted to counteract vibratory aerodynamic loads.
The concept of using AFCs for integral rotor blade actuation was first demonstrated in 1998.
A first generation scaled blade was developed by MIT and Boeing Helicopters and was hover
tested. This blade experience several AFC actuator failures, either through loss of electrical
connection or through internal short circuits. The blade also did not meet peak actuation
expectations. This thesis advances the integral actuation technology to avoid the problems
experienced in the first generation blade. Manufacturability improvements are made through
redesign of blade components associated with the active material. These include the blade power
bus, the AFC actuator pack leadouts, and the electrical connections to the AFCs. Material
level work focuses on the mechanical and electrical behavior of embedded AFCs. It was shown
that the mechanical behavior of the AFCs is non-linear and voltage dependent. The dielectric
properties of the AFCs are studied and shown to be strongly voltage dependent.
A second generation integrally actuated blade is built in conjunction with Boeing Helicopters. Bench testing results of this blade are presented. The blade performed as expected,
yielding a 2.3*/m twist rate through the span of the blade at 3000Vpp, OVDC drive voltage.
Blade heating and current draw were observed to be significant, but in line with expectations
based on testing of individual AFCs.
Thesis Supervisor: Nesbitt Hagood IV
Title: Associate Professor, Department of Aeronautics and Astronautics
Mechanical Engineering Faculty Reader: Seth Lloyd
Title: Assistant Professor, Department of Mechanical Engineering
3
4
Acknowledgements
It is with heartfelt gratitude that I dedicate this thesis to all those who have played a role in its
successful completion.
I would like to thank my family and my friends for their support and
encouragement through the past two and a half years. Lots of hugs and kisses go to Amanda
for sticking by my side through all the hard times - I promise I'll be happier from now on.
Funding for this research was provided by the Defense Advanced Research Projects Agency,
under contract MDAA72-98-3-0001,
and was monitored by Ephrahim Garcia of the DARPA
Defense Sciences Office.
5
6
Contents
1
2
3
19
Introduction
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
19
1.1
Motivation
1.2
Active Fiber Composites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22
1.3
Objectives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
1.4
Approach
1.5
Summary of Document . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
State of Integral Actuation Technology
29
2.1
Overview of Blade Mounted Actuator Technologies . . . . . . . . . . . . . . . . . 29
2.2
Integrally Actuated Blades
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
2.2.1
Actuator
2.2.2
Blade Scaling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32
2.2.3
MIT/NASA Langley Integral Blade
2.2.4
MIT/Boeing Integral Blade . . . . . . . . . . . . . . . . . . . . . . . . . . 33
. . . . . . . . . . . . . . . . . . . . . 32
41
Active Fiber Composite Testing
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41
3.1
Motivation
3.2
Approach
3.3
AFC Properties Review
3.4
Modeling Concept
3.5
Coupon Lamination Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46
3.5.1
Test Samples
3.5.2
Test Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46
7
3.6
3.7
4
3.5.3
Lamination Procedure.....
3.5.4
Results
3.5.5
Summary ..............
48
. . . . . . . . . . . . .
Blade Pack Lamination Testing
50
50
52
.
3.6.1
Motivation
. . . . . . . . . . .
52
3.6.2
Test Procedure . . . . . . . . .
52
3.6.3
Results
. . . . . . . . . . . . .
54
3.6.4
Analysis . . . . . . . . . . . . .
54
3.6.5
Summary . . . . . . . . . . . .
57
Blade Pack Heating and Current Draw Testing
57
3.7.1
Motivation
57
3.7.2
Test Samples
. . . . . . . . . . . . . . . . . . . . . . . . .
58
3.7.3
Test Procedure . . . . . . . . . . . . . . . . . . . . . . . .
58
3.7.4
Heating Results and Analysis
. . . . . . . . . . . . . . .
60
3.7.5
Current Draw Results.....
. . . . . . . . . . . . . . .
61
3.7.6
Power Loss . . . . . . . . . . . . . . . . . . . . . . . . . .
63
3.7.7
Blade Current Requirements and Performance Envelope .
74
3.7.8
Summary . . . . . . . . . . . . . . . . . . . . . . . . . . .
75
. . . . . . . . . . .
Design and Manufacturing of Internal High Voltage Connections
77
4.1
Motivation
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
77
4.2
Approach
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
78
4.3
Actuator Pack Leadout Design
. . . . . . . . . . . . . . . . . . . . .
79
4.4
Flexible Circuit Design . . . . . . . . . . . . . . . . . . . . . . . . . .
81
4.4.1
Motivation for Using a Flexible Circuit
. . . . . . . . . . . .
81
4.4.2
Flexible Circuit Design
. . . . . . . . . . . . . . . . . . . . .
82
Blade Manufacturing Procedure . . . . . . . . . . . . . . . . . . . . .
83
4.5.1
Blade Section 4.0 Overview . . . . . . . . . . . . . . . . . . .
83
4.5.2
Spar Layup . . . . . . . . . . . . . . . . . . . . . . . . . . . .
85
4.5.3
Electrode Tab Protection
. . . . . . . . . . . . . . . . . . . .
86
4.5.4
Spar Cure . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
86
4.5
8
4.5.5
5
4.6
Blade Section 4.0 Manufacturing Results
4.7
Actuation Testing Setup and Procedure
4.8
Blade Section 4.0 Testing
4.9
Summary
. 94
94
. . . . . . . . .
98
. . . . . . . . . . . . . . . . . .
. 99
101
AMR Blade Design and Testing
5.1
6
. 89
Fairing Assembly . . . . . . . . . .
. . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . 102
5.1.1
Blade Geometry and Layup . . . .
. . . . . . . . . . . . . . . . . . . 102
5.1.2
Actuator Packs . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . 102
5.1.3
Internal Flexible Circuit Details
. . . . . . . . . . . . . . . . . . . 10 5
5.1.4
External Circuit Design . . . . . .
. . . . . . . . . . . . . . . . . . . 106
5.1.5
Power Distribution Board.....
. . . . . . . . . . . . . . . . . . . 1 10
5.1.6
Options in Case of Flexible Circuit
AMR Design
Failu re
. . . . . . . . . . . . . . . . . 110
. . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . 111
. .
. . . . . . . . . . . . . . . . . . . 1 12
5.3.1
Testing Setup . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . . . 1 16
5.3.2
Actuation Testing Results.....
. . . . . . . . . . . . . . . . . . . 1 17
5.3.3
Blade Mechanical Properties
5.3.4
Blade Current Draw and Heating
. . . . . . . . . . . . . . . . . . . 124
5.3.5
Actuator Pack Failure History
.
. . . . . . . . . . . . . . . . . . . 129
5.3.6
Destructive Testing . . . . . . . . .
5.3.7
Windtunnel Testing Recommendations . . . . . . . . . . . . . . . . . . . . 134
5.2
Manufacturing
5.3
Bench Testing of AMR Process Blade
. . .
. . . . . . . . . . . . . . . . . . . 121
. . . . . . . . . . . . . . . . . . . 132
Conclusions and Recommendations
137
6.1
Summary and Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
137
6.2
Recommendations for Future Work . . . . . . . . . . . . . . . . . . . . . . . .
139
A Coupon Testing Data
149
B Blade Section 4.0 Supplement
153
9
C AMR Design and Testing Supplement
C.1
159
Blade Actuator Pack Data ...............................
C.2 Bending Stiffness . . ..
C.3 Blade Dynamics
. . . ....
.........
......
159
..
. . . . . . . . . . ...
....................................
. ..
. 159
159
C.4 Blade Crossectional Pictures ....................................
161
C.5 AM R Flexible Circuit
161
.................................
D Flexible Circuit Connectors
173
E Piezoelectric Ceramic Data
179
F Core Void Elimination
181
F.1 M otivation
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 181
F.2
Hypotheses of void formation ......
F.3
Test procedure
F .4
R esults . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 185
F .5
A nalysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 186
F.6
Recommendations
........
.............................
.....................................
182
183
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 188
F.7 Voids Formation Testing Details
. . . . . . . . . . . . . . . . . . . . . . . . . . . 189
10
List of Figures
1-1
Active blade concept . . . . . . . . . . . . . . . . . . . . . . . . .
23
1-2
Active fiber composite concept
. . . . . . . . . . . . . . . . . . .
24
1-3
Electric field in AFC fiber . . . . . . . . . . . . . . . . . . . . . .
24
2-1
Geometry and layup of the MIT/Boeing integrally actuated rotor blade
2-2
Integral blade on the MIT hover test stand
2-3
CH-47D blade section and full blade twist rates . . . . . . . . . . . . . . . . . . . 35
2-4
First generation blade hover data . . . . . . . . . . . . . . . . . . . . . . . . . . .
2-5
Rodgers blade connections . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38
2-6
Damage map of first generation blade
3-1
Simple constraint model . . . . . . . . . . . .
. . . . . . . . . . 44
3-2
Actuator and constraint stress-strain plots . .
. . . . . . . . . . 44
3-3
AFC standard test coupon . . . . . . . . . . .
. . . . . . . . . . 47
3-4
Setup for coupon lamination in the hotpress .
. . . . . . . . . . 49
3-5
Actuation testing direction
. . . . . . . . . .
. . . . . . . . . . 53
3-6
Blade pack lamination data . . . . . . . . . .
. . . . . . . . . . 55
3-7
Reduction factors for high and low free strain AFC packs
. . . . . . . . . . 56
3-8
Time histories of heating of unlaminated blade packs . . .
. . . . . . . . . . 60
3-9
Effect of DC offset on blade pack heating
. . . . . . . . .
. . . . . . . . . . 61
3-10 Current draw of unlaminated packs U2 and U3 . . . . . .
. . . . . . . . . . 62
3-11 Current draw of laminated packs 1H and 1L . . . . . . . .
. . . . . . . . . . 62
. . . . . . . . . . . . . .
. . . . . . . . . . 63
3-12 Peak current draw vs. frequency
11
. . . . . 34
. . . . . . . . . . . . . . . . . . . . . 34
36
. . . . . . . . . . . . . . . . . . . . . . . . 38
3-13 Effect on DC offset on current draw
. . . . . .
. . . . . . . . . . . . . . . . 64
3-14 Equivalent circuit model for the AFCs . . . . .
. . . . . . . . . . . . . . . . 65
3-15 Fit of R-C model to current-voltage data . . . .
. . . . . . . . . . . . . . . . 66
3-16 Capacitance and resistance for laminated packs
. . . . . . . . . . . . . . . . 68
3-17 Dissipated AFC current . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . 69
3-18 Power draw data for laminated packs . . . . . .
. . . . . . . . . . . . . . . . 70
3-19 Tan6 data for laminated packs
. . . . . . . . . . . . . . . . 71
. . . . . . . . .
3-20 Performance envelope for AMR blade
. . . . .
. . . . . . . . . . . . . . . . 74
4-1
Active materials rotor blade showing the routing of the fle xible circuit
4-2
Typical blade actuator pack . . . . . . . . . . . . . . . . .
. . . . . . . . . . . 80
4-3
Electrode leadout tab detail . . . . . . . . . . . . . . . . .
. . . . . . . . . . . 80
4-4
Blowup of flexible circut . . . . . . . . . . . . . . . . . . .
. . . . . . . . . . . 82
4-5
Blade section 4.0 drawing
. . . . . . . . . . . . . . . . . .
. . . . . . . . 84
4-6
Blade section 4.0 spar layup . . . . . . . . . . . . . . . . .
. . . . . . . . 85
4-7
Protection of electrode tab during spar cure . . . . . . . .
. . . . . . . . 86
4-8
Hardback manufacturing . . . . . . . . . . . . . . . . . . .
. . . . . . . . 87
4-9
Hardback mounted in blade mold . . . . . . . . . . . . . .
. . . . . . . . 87
.............
4-10 Section 4 spar.............................
. . . . . . 78
89
4-11 AMR blade fairing assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90
4-12 Flexible circuit routing on fairing foam core . . . . . . . . . . . . . . . . . . . . . 91
4-13 Blade section 4.0 with connections to flexible circuit made . . . . . . . . . . . . . 91
4-14 Schematic of flexible circuit transition from web to fairing . . . . . . . . . . . . . 92
4-15 Connection area filler material during fairing cure
. . . . . . . . . . . . . . . . . 92
4-16 Completed blade section 4.0 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95
4-17 Active area of completed blade section 4.0 . . . . . . . . . . . . . . . . . . . . . . 97
4-18 Twist measurement schematic . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97
4-19 Blade section 4.0 actuation
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98
. . . . . . . . . . . . . . . . . . . . . . . . . 103
5-1
Active Material Rotor (AMR) blade
5-2
AM R blade layup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103
12
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104
5-3
AMR AFC actuator pack
5-4
Internal flexible circuit . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10 5
5-5
AMR blade crossection showing the flexible circuit . . . . . . . . . . . . . . . . . 106
5-6
Oversized solder pad cutout in flexible circuit . . . . . . . . . . . . . . . . . . . . 10 7
5-7
Rotor hub with flex circuit routing . . . . . . . . . . . . . . . . . . . . . . . . . . 108
5-8
AMR external flexible circuit
5-9
Power distribution board . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108
5-10 Power distribution board with flexible circuits . . . . . . . . . . . . . . . . . . . . 109
5-11 Process blade actuator packs allocation
. . . . . . . . . . . . . . . . . . . . . . . 111
5-12 Completed AMR blade . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 113
5-13 AMR blade root . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114
5-14 AMR blade midspan detail
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
5-15 AMR blade tip detail . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115
5-16 Actuation testing setup
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118
5-17 AMR blade clamped at root . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118
5-18 Actuation of the process blade at 1Hz . . . . . . . . . . . . . . . . . . . . . . . . 120
5-19 AMR blade actuation data at several frequencies. . . . . . . . . . . . . . . . . . . 120
5-20 AMR blade twist versus span data . . . . . . . . . . . . . . . . . . . . . . . . . . 12 1
5-21 Depoling of blade AFCs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 122
5-22 Effect of a DC offset on the depoling behavior of the blade AFCs . . . . . . . . . 122
5-23 Torsional stiffness testing setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123
5-24 Blade current draw . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124
5-25 Frequency dependence of blade heating . . . . . . . . . . . . . . . . . . . . . . . . 126
5-26 Voltage dependence of blade heating . . . . . . . . . . . . . . . . . . . . . . . . . 126
5-27 DC offset dependence of blade heating . . . . . . . . . . . . . . . . . . . . . . . . 127
5-28 Effect of DC offset on current draw . . . . . . . . . . . . . . . . . . . . . . . . . . 127
5-29 Blade capacitance and tan
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 128
5-30 Actuator pack failure schematic . . . . . . . . . . . . . . . . . . . . . . . . . . . . 130
5-31 Blade crossection . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133
5-32 Crossection of AFC failure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133
13
5-33 Crossection of connections . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13 3
5-34 Crossection of flexible circuit transition
. . . . . . . . . . . . . . . . . . . . . . . 13 3
B-1 Section 4.0 parts list . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 5 4
B-2 Section 4.0 drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 5
B-3 Section 4.0 flexible circuit design . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 5 6
B-4 Drawing of outboard tip fitting . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1 5 7
B-5 Section 4.0 electrode pattern
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 8
C-1 Average bending stiffness of the AMR blade . . . . . . . . . . . . . . . . . . . . . 160
C-2 Transfer function from drive voltage to twist rate . . . . . . . . . . . . . . . . . . 162
C-3 Transfer function from drive voltage to blade curvature
. . . . . . . . . . . . . . 162
C-4 AMR blade average spectrum in twist mode . . . . . . . . . . . . . . . . . . . . . 163
C-5 AMR blade average spectrum in bending mode . . . . . . . . . . . . . . . . . . . 163
C-6 Second bending mode of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . 164
C-7 Second bending mode of the AMR blade.
. . . . . . . . . . . . . . . . . . . . . . 164
C-8 Third bending mode of the AMR blade
. . . . . . . . . . . . . . . . . . . . . . . 165
C-9 Third bending mode of the AMR blade
. . . . . . . . . . . . . . . . . . . . . . . 165
C-10 Fourth bending mode of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . 166
C-11 Fourth bending mode of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . 166
C-12 First torsion mode of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . . . 167
C-13 First torsion mode of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . . . 167
C-14 Crossections of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . . . . . . 168
C-15 Crossections of the AMR blade . . . . . . . . . . . . . . . . . . . . . . . . . . . . 169
C-16 AMR flexible circuit drawing
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . 170
C-17 Location of the solder pads in the AMR flexible circuit . . . . . . . . . . . . . . . 171
D-1 Flexible circuit flareout trimming . . . . . . . . . . . . . . . . . . . . . . . . . . . 175
D-2 Flexible circuit connector crimping . . . . . . . . . . . . . . . . . . . . . . . . . . 176
D-3 Flexible circuit crimp connector soldering . . . . . . . . . . . . . . . . . . . . . . 176
D-4 Pin insertion into crimp connectors . . . . . . . . . . . . . . . . . . . . . . . . . . 177
D-5 Flex circuit connector mold . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177
14
D-6
Flexible circuit connector potting . . . . . . . . . . . . . . . . . . . . . . . . . . . 178
E-1 eT and tan 6 information on PZT-5A, PZT-4, and PZT-8 . . . . . . . . . . . . . 180
F-i
First generation blade damage map . . . . . .
. . . . . . . . . . . . . . . . . 182
F-2 Crossection of first generation blade at BS19
. . . . . . . . . . . . . . . . . 182
F-3 Voids testing spar assembly . . . . . . . . . .
. . . . . . . . . . . . . . . . . 184
F-4 Voids testing spar layup . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 184
F-5 Spar crossectioning . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 185
F-6 Void locations in test spar . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 186
. . . . . . .
. . . . . . . . . . . . . . . . . 187
. . . . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 188
F-9 Voids comparison . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 189
. . . . . . . . . . . . . . . . .
. . . . . . . . . . . . . . . . . 190
F-7 Picture of voids in the test piece
F-8 Void detail
F-10 Moisture voids
F-11 Foam block 1: Crossections 1-4, outboard view. . . . . . . . . . . . . . . . . . . . 190
F-12 Foam block 1: Crossections 1-4, inboard view. . . . . . . . . . . . . . . . . . . . . 19 1
F-13 Foam block 2: Crossections 5-9, outboard view. . . . . . . . . . . . . . . . . . . . 19 1
F-14 Foam block 2: Crossections 5-9, inboard view. . . . . . . . . . . . . . . . . . . . . 19 1
F-15 Foam block 3: Crossections 10-14, outboard view. . . . . . . . . . . . . . . . . . . 192
F-16 Foam block 3: Crossections 10-14, inboard view.
. . . . . . . . . . . . . . . . . 192
F-17 Foam block 4: Crossections 15-20, outboard view. . . . . . . . . . . . . . . . . . . 192
F-18 Foam block 4: Crossections 15-20, inboard view.
. . . . . . . . . . . . . . . . . 193
F-19 Foam block 5: Crossections 21-25, outboard view. . . . . . . . . . . . . . . . . . . 193
F-20 Foam block 5: Crossections 21-25, inboard view.
. . . . . . . . . . . . . . . . . 193
F-21 Foam block 6: Crossections 26-29, outboard view. . . . . . . . . . . . . . . . . . . 194
F-22 Foam block 6: Crossections 26-29, inboard view.
15
. . . . . . . . . . . . . . . . . 194
16
List of Tables
3.1
Nominal low stress AFC and E-glass properties . . . . . . . . . . . . . . . . . . . 43
3.2
Coupon lamination testing test matrix . . . . . . . . . . . . . . . . . . . . . . . . 47
3.3
Summary of coupon testing, control and heat-only specimens
. . . . . . . . . . . 51
3.4
Summary of coupon lamination testing, cure pressure results
. . . . . . . . . . . 51
3.5
Blade pack lamination testing pack description
3.6
Summary of blade pack lamination testing . . . . . . . . . . . . . . . . . . . . . . 54
3.7
Heating and current draw testing pack description
5.1
Failure history of process blade actuator packs. . .
130
A.1
Results of coupon testing
. . . . . . . . . . . . . .
150
. . . . . . . . .
151
. . . . . . . . . .
152
C.1 Actuation data for process blade packs . . . . . . .
160
Piezoelectric ceramic properties . . . . . . . . . . .
179
A.2 Capacitance data for test coupons
A.3 Resistance data for test coupons
E.1
17
. . . . . . . . . . . . . . . . . . . 53
. . . . . . . . . . . . . . . . . 59
18
Chapter 1
Introduction
Integrally actuated helicopter rotor blades are being developed by MIT and Boeing Helicopters
[Rodgers, 1998; Derham, 1996; Schmidt, 2000]. Piezoelectric fibers, in the form of Active Fiber
Composites (AFCs), are incorporated into the skin of the blades to enable active deformation
of the structure.
A prototype integrally actuated blade was developed in 1998 by Rodgers.
This blade demonstrated that the concept of embedding AFCs in a helicopter blade was feasible, but fell short of expected performance levels [Rodgers, 1998].
This thesis will focus on
the further development of integral actuation technology for helicopter rotor craft structures.
This development work will focus on several problems that were encountered in the Rodgers
blade, and will involve improving the blade manufacturing procedure and acquiring a better
understanding of the behavior of the active material embedded in the blade.
1.1
Motivation
Vibration has historically been a great concern of helicopter manufacturers, pilots, and passengers.
Much work has been done to reduce vibration over the past 50 years, and as a result
vibratory loads in helicopters have been steadily decreasing.
continued effort to dampen vibration in helicopters:
Two criteria have driven the
human factors, such as ride comfort and
cabin noise, and maintainability and functional reliability [Reichert, 1980].
Non-uniform airflow through the rotor in forward flight causes periodic vibratory aerodynamic forces on the rotor. These vibratory loads usually exhibit themselves as n-per revolution
19
(n-PRev) excitation forces on the hub, where n is the number of blades in the rotor. Harmonics of this frequency are also excited, leading to a broader spectrum of vibratory loads [Braun,
1984].
Atmospheric turbulence, retreating blade stall, blade vortex interaction, and blade fuselage
interference are contributing factors in the generation of rotor vibration [Ham, 1987].
These
sources of vibration apply moments and forces to the hub of the rotor, and vibration is transferred to the fuselage through the drive shaft and motor.
Passive damping has been used for many years to attempt to reduce the vibratory loads
transmitted from the rotor to the fuselage.
Passive damping work has focused on altering
the dynamics of the rotor and of the fuselage to cancel out vibration at certain frequencies.
Pendulum absorbers can be mounted at the root of the blade to create a node at a given
frequency.
1974].
50% reduction in vibratory loads have been achieved with such systems [Amer,
Isolation systems can also be implemented between the fuselage and the drive unit and
rotor. These systems alter the dynamics of the vehicle to create a point of zero vibration at a
given frequency [Desjardins, 1978].
The disadvantage of using passive dampers to reduce vibration in helicopters is that passive
dampers are generally tuned to a specific frequency.
Since vibration usually occurs over a
wider spectrum of frequencies, including many harmonics of the primary rotor vibration, passive
vibration isolation or damping can only have limited success.
The use of active vibration control can be used to dampen vibration over a larger range
of frequencies.
Two main approaches are taken towards active vibration control:
Harmonic Control (HHC) and Individual Blade Control (IBC).
Higher
HHC involves exciting all
blades with the same signal to attempt to counteract vibratory airloads.
This actuation
method is usually implemented through oscillatory swash plate motion to cause periodic pitch
control [Shaw, 1981].
frequency.
Vibration reduction is mostly concerned with harmonics of the rotor
Due to rotor symmetry, only these harmonics are transmitted to the fuselage.
HHC is therefore well suited for vibration control. However, for noise control, a much wider
spectrum must be considered since structural filtering does not occur in each of the individual
blades [Prechtl, 2000].
Therefore,
HHC is not a practical method for simultaneous noise
reduction and structural vibration reduction.
20
IBC involves controlling each blade separately.
Each blade has an actuator mechanism
and an individual feedback loop to allow control of each blade independently.
This provides
more degrees of freedom of the rotor than HHC allows, leading to more effective reduction of
Individual blade control can be implemented at the hub through
vibration [Jacklin, 1995].
replacing the rotating pitch links with hydraulic actuators.
These actuators can be used to
apply small pitching motions at high frequencies independently to each blade [Swanson, 1994].
Alternatively, IBC can be implemented through blade mounted actuators.
This imple-
mentation allows vibration reduction at the source of the disturbance [Reichert, 1980].
Blade
mounted actuator technology has seen rapid growth due to the development of smart actuation
technology. The use of smart materials, such as piezoelectric ceramics and shape memory alloys (SMA), allows direct conversion of electrical energy to mechanical energy. This eliminates
the need for complex mechanisms and hydraulic systems to actuate the blades at high frequencies.
Induced strain actuators have large force and energy capabilities, but they have very
small strokes. Therefore, in most cases, displacement amplification mechanisms must be used
to create usable actuator deflections [Giurgiutiu, 2000].
A great advantage to blade mounted
actuators is that they are not critical flight elements. A failure of an actuator is therefore not a
catastrophic event since it does not compromise the robustness of the vehicle, as, for example,
a pitch link actuator failure would [Prechtl, 2000].
A direct approach to counteracting vibrational airloads on the rotor is distributed twist
actuation of the blades [DuPlessis, 1995]. See Figure 1-1. Embedding active material directly in
the skin of the blades can eliminate the need for stroke amplification mechanisms and transform
the entire blade structure into an actuator. Distributed actuation has several advantages. Since
there are no moving parts, such as displacement amplification mechanisms or flap assemblies,
maintenance is simplified.
The lack of mechanisms and flaps in the blade allows actuation
without adding drag due to extra aerodynamic surfaces.
redundancy in actuation.
packs.
Distributed actuation also means
The active material is incorporated into the blade skin in several
If one or more packs fail, the remaining packs can continue to actuate the blade,
although at a reduced level.
Active Fiber Composites (AFCs) have been developed for the purpose of embedding in
composite structures.
These actuators provide anisotropic, in plane actuation, and when
21
oriented at a 45 angle to the blade span can be used to create a distributed twisting moment
along the span of the blade.
This moment causes the blade to twist, changing the lift of the
blade, enabling implementation of IBC.
A blade was developed by Rodgers in 1998 that implemented this technology with mixed
success.
The blade experienced several actuator failures due to manufacturing and opera-
tional stresses.
These failures were partly caused by actuator material immaturity and blade
manufacturing difficulties.
Due to actuator failures, the drive voltage of the Rodgers blade
was limited to 2000Vpp, instead of the full design voltage of 4000Vpp.
The lowered voltage,
along with failure of approximately half the actuator packs prior to hover testing, limited the
blade actuation level to approximately 25% of the design level. The blade did however demonstrate that a structurally sound active blade could be built and actuated in a realistic hover
environment [Rodgers, 1998]. More information on actuator failures is provided in Chapter 2.
1.2
Active Fiber Composites
Active Fiber Composites use piezoelectric fibers for in-plane actuation. Piezoelectric materials
exhibit coupling between electrical and mechanical behavior.
Mechanical displacement will
cause charge to flow through the material, and the application of an electric field will cause
mechanical deformation. Extensive information on piezoelectric ceramics have been published,
and will not be discussed here [Bent, 1997; Ghandi, 1998; Jaffe, 1971; Berlincourt, 19XX].
Most piezoelectric actuators are based on piezoelectric wafers, either as single wafers or as
stacks of wafers. AFCs, on the other hand, consist of piezoelectric fibers, 0.005" to 0.01" thick,
embedded in an epoxy matrix.
low strains.
The ceramic fibers are brittle and tend to crack at relatively
Surrounding the fibers with an epoxy matrix significantly improves strength and
toughness properties of the actuator. The matrix reinforces the fibers, allowing the actuator to
withstand higher strains than the individual fibers could withstand. The epoxy matrix provides
a path for load transfer around fiber cracks. This keeps a crack on one fiber from propagating
to other fibers and causing extensive damage to the AFC. This presents an improvement over
actuators based on ceramic wafers, which are prone to macroscopic and catastrophic damage
resulting from cracks in the ceramic [Jones, 1999].
22
.
.
.
..
. .. .................
............................
.....
........
Passive Plies
Active Plies
Aerodynamic
Fairing
Weights
Sn
ar-
Web
Foam Core
Figure 1-1: Active blade concept. Active plies are embedded among the passive plies. In
plane actuation of active plies oriented at +/- 450 induces a torsional moment along the span
of the blade, twisting the blade.
AFCs are equipped with an interdigitated electrode pattern. See Figure 1-2. This electrode
pattern orients the electric field along the fiber, taking advantage of the primary piezoelectric
effect, the expansion of the ceramic in the direction of the applied field.
See Figure 1-3.
Excitation of the actuator therefore causes in-plane actuation in the direction of the fibers.
This actuation mode is ideal for use in adaptive composite structures.
These packs are of a given geometry, which
The AFCs are manufactured in "packs".
is application dependent.
For example, the AMR blade, which will be discussed in detail
in Chapter 5, incorporates 24 individual packs into the blade.
independent, so if one pack fails, it will not affect the other packs.
23
These packs are electrically
.............................
Figure 1-2: Active fiber composites (AFCs) consist of piezoelectric fibers embedded in an epoxy
matrix, surrounded by an interdigitated electrode pattern. The electrode pattern orients the
electric field along the fiber, taking advantage of the primary piezoelectric effect of the fibers.
See Figure 1-3.
A-*
Electrode Fingers
1(
Figure 1-3: Schematic representation of the electric field in an individual fiber. Interdigitated
electrodes orient the electric field along the fiber. This allows the actuator to take advantage
of the primary piezoelectric effect, the extension in the direction of the applied field. In plane
actuation of the AFC is achieved using this actuation method.
24
1.3
Objectives
The main objective of this thesis is to further develop and improve integral actuation technology
using AFCs and to incorporate this technology into a composite rotor blade structure.
Two
main areas will be addressed:
" Actuator understanding
" Manufacturing improvements
Improving and simplifying the manufacturing procedure of the active blades is important
to creating more robust and better performing helicopter blades.
embedded in the skin of the blade, they can not be replaced.
actuator failures be minimized.
during operation.
Since the actuators are
Therefore, it is important that
Actuator failures can occur either during manufacturing or
Improving the design of the integral actuation system, including AFCs,
actuator leadouts, and high voltage lines, will reduce the number of failures that occur during
manufacturing.
It will also create a more robust blade, minimizing the number of actuator
failures that occur during operation.
In designing integrally actuated blades that meet actuation and structural requirements, it
is necessary to be able to accurately predict the behavior of the actuators when embedded in
the blade.
Historically, modeling of the blades has been difficult because of the unpredictable
A better understanding of the high field actuation non-linearities of the
nature of the AFCs.
AFCs is necessary to improve the accuracy of blade modeling.
Determining the electrical behavior of the AFCs embedded in the blades will provide valuable information necessary for practical application of the AFCs.
Voltage and frequency
dependence of actuator current draw will be studied. This information will help determine the
amount of current necessary to drive the blades at a given voltage and frequency, and will help
guide the selection of amplifiers to be used in windtunnel testing of the blades.
1.4
Approach
Lessons learned in the Rodgers blade will be applied to develop the new design of the integral
actuation system.
The manufacturing procedure used in the Rodgers blade will be studied
25
and improvements made.
The main improvement will involve redesign of the power bus, the
set of power leads that bring power to the actuators.
to the actuators will be improved and simplified.
Manufacturability of the connections
Hypotheses for actuator failures presented
by Rodgers will be studied and appropriate changes will be applied to the new design of the
integral actuation system [Rodgers, 19981.
Actuator understanding will be developed through testing of individual AFC packs. Lamination testing will be performed to determine the effects of constraint on the actuation authority
of the actuators.
the AFCs.
An equivalent circuit model will be developed to study the current draw of
This model will provide information on the voltage and frequency dependence of
AFC capacitance and dielectric loss.
Finally, an active blade will be designed and bench tested.
actuation testing, mechanical testing, and electrical testing.
Bench testing will include
The results will give further
information on mechanical and electrical AFC behavior, and will show how the behavior of
individual AFCs correlates with the behavior of the blade structure.
1.5
Summary of Document
Chapter 2 will describe the current state of integral actuation technology. The first generation
integrally actuated blade will be presented.
Actuation capability and a summary of hover
testing will be given. The problems associated with the first generation blade and their impact
on the performance of the blade will then be discussed. Hypotheses for the numerous actuator
failures that occurred in this blade will be discussed, focusing mainly on blade manufacturing.
Chapter 3 will present the actuator characterization work performed to better understand
the behavior of the embedded actuators. AFC coupon and blade packs are tested to determine
the effects of lamination on the AFCs. Significant variations in laminated performance are seen
in the packs.
Recommendations on reduction factors to be used in modeling the integrally
actuated blades will be given.
Extensive testing is performed on blade packs to understand
the electrical behavior of the AFCs.
Analysis of the high current draw and heating of the
AFCs will be presented. Drive voltage and frequency dependence of actuator pack capacitance
and dielectric loss factor will be quantified using a simple equivalent electrical circuit.
26
It will
be shown that both the pack capacitance and dielectric loss factor increase significantly as the
drive voltage is incresed.
Recommendations for testing of the integral blades will be given
based on the data presented in this chapter.
Chapter 4 will describe the work done to improve the manufacturing of the blade. Significant
changes are made to the power bus design and to the connections to the actuators. The power
bus and connections to the actuators is moved from the web of the blade to the surface of the
fairing. The design of the actuator leadouts, the flexible circuit power bus, and the connections
between the two will be presented. To test the new design, manufacturing studies are conducted.
These manufacturing studies include building a pair of short blade sections incorporating the
new design.
The new design was shown to be a viable design for incorporation into a full
model-scale active blade.
Chapter 5 will present the Active Material Rotor (AMR) blade being developed for wind
tunnel testing at Boeing, Philadelphia.
The implementation of the design changes discussed
in Chapter 4 will be the main focus of the first part of this chapter.
chapter will focus on the testing of the blade.
electrical testing, and mechanical testing.
The second part of the
This testing will include actuation testing,
The blade achieved a 2.30 /m twist rate, compared
to a 2.50 /m model prediction. Analysis of actuator failures will be conducted.
Mechanical
testing will give the torsional and bending stiffnesses of the blade, and show correlation to
models.
The current draw was greater than expected, as was the heating, but both were
in-line with expectations based on individual pack testing. Electrical data will also include the
blade capacitance and loss factors as a function of voltage.
These results will be compared to
the results of individual actuator pack testing.
Chapter 6 will conclude the thesis by summarizing the accomplishments of the work and
providing recommendations for changes to the actuator design, blade design, and manufacturing
procedure.
Appendices are included to provide blade manufacturing information and actuation
data that was not included in the body of the thesis.
27
28
Chapter 2
State of Integral Actuation
Technology
This chapter will discuss the current state of blade mounted actuation technology, focusing
on integral actuation using Active Fiber Composites. Past work on the actuators and on the
blades will be described.
2.1
Other blade mounted actuator technologies will be discussed briefly.
Overview of Blade Mounted Actuator Technologies
Smart material actuators have helped overcome the size, weight, and complexity problems that
have limited the incorporation of hydraulic and electromechanical actuators into helicopter
blades (Straub, 1996].
This section will provide a brief overview of various smart material
approaches to blade mounted actuation.
A discretely actuated blade has been developed at MIT by Prechtl and Hall, in conjunction
with Boeing Helicopters. This actuation scheme is based on using a piezoelectric stack actuator
with a displacement amplification mechanism to drive a trailing edge flap. The blade has had
success in vibration reduction studies, virtually eliminating vibration at several rotor harmonics
simultaneously [Prechtl, 2000b].
Prahlad and Chopra at the University of Maryland are developing a rotorblade that incorporates Shape Memory Alloy (SMA) into the structure.
The change in Young's Modulus
of the material as a function of temperature allows the natural frequencies of the blade to be
29
changed.
The dynamic properties of the rotorblade can therefore be tuned to varying blade
environments [Prahlad, 2000].
variable camber blades.
The actuation capability of SMA technology is used to create
By embedding SMA in the blade skin, the camber of the blade can be
actively changed to alter the aerodynamic properties of the airfoil [Roglin, 1994].
Spangler and Hall at MIT developed a flap actuator based on a piezoelectric bimorph.
tip displacement
The
of the bimorph actuator is used to drive a trailing edge flap through an
amplification mechanism.
This actuator is simple, lightweight, and compact, and is therefore
suited well for integration into a helicopter blade [Spangler, 1990].
Walz and Chopra, at the
University of Maryland, have developed a similar actuator [Walz, 1994].
An actuator taking advantage of induced shear properties of piezoceramics is being developed by Centolanza and Smith at Pennsylvania State University.
The actuator incorporates
electrodes in the tube that allow the shear piezoelectric effect to be excited, twisting the tube
along the length. The piezoceramic tube is used to drive a trailing edge flap through a coupling
mechanism [Centolanza, 2000].
A full scale discrete actuator is being incorporated into an MD-900 blade by Boeing Helicopters, Mesa.
This blade incorporates a piezo stack actuator for high frequency actuation,
and an SMA actuator for low frequency trim tab actuation [Straub, 2000].
Integrally actuated blades based on piezoelectric ceramic actuators have been developed
by MIT, Boeing, NASA, and the University of Maryland.
The University of Maryland blade
incorporated piezoelectric wafers within the skin of a rotor blade.
into strips and were oriented at 450 to the blade span.
The wafers were sliced
This program had moderate success,
achieving only small deflection angles of the blade [Chen, 1994].
With the development of AFCs, integral actuation using piezoelectric materials has become
a more viable actuation technology.
AFCs are more easily incorporated in the composite
structures, due to their small thickness, their flexible nature, and their in-plane actuation
capability.
30
2.2
2.2.1
Integrally Actuated Blades
Actuator
AFC actuators have been under steady development over the past decade. The material system
has become more reliable and manufacturability has been improved [Strock, 1999]. The cost of
the actuator system has also decreased significantly.
Until 1998, AFCs had been manufactured using copper-Kapton electrodes. These electrodes
consisted of an etched interdigitated copper electrode pattern on a Kapton substrate [Rodgers,
1998]. Due to cost and manufacturing difficulties, the copper-Kapton electrodes were abandoned
in favor of screen printed silver ink electrodes. These electrodes are screen printed onto a Kapton
substrate using conductive ink. Because the screen printed electrodes are more malleable, they
conform better to the shape of the fibers, providing a greater contact area between the electrodes
and the fibers. This increases the actuation authority by introducing more of the electric field
into the fiber [Wickramasinghe, 2000].
The fibers that were used in the AFCs until late 1999 were 0.005" diameter fibers. These
fibers were manufactured by CeraNova Corporation'. AFCs manufactured with 0.005" diameter
fibers had a total thickness of between 0.008" and 0.009".
For comparison, the passive plies
used in the blade are all under 0.010" in thickness. The AFCs are therefore thin enough to be
easily incorporated into the skin of the blades.
To improve manufacturability and scalability to full size, 10 mil fibers were introduced in the
fall of 1999 as a possible alternative to the 5 mil fibers. AFCs manufactured with 0.010" fibers
have a total thickness of 0.013". Properties of the 0.005" and 0.010" fiber AFCs are presented in
Chapter 3. There are several advantages to using 0.010" fibers. Since one of the determining
factors of the actuation authority is the volume of active material, only one quarter as many
0.010" fibers as 0.005" fibers are necessary to create a given actuation authority.
Since the
manufacturing of fibers, and not the bulk material or size of the fibers, is the most significant
factor driving the cost of the AFCs, using larger diameter fibers translates to large cost savings.
The 0.010" fibers, being straighter than the 0.005" fibers, can also help achieve line fractions, or
fiber density, higher than in the 0.005" fiber AFCs. Line fractions as high as 90% are achievable
'CeraNova Corporation, Franklin, MA.
31
with the larger fibers, further increasing actuation authority [Pizzocchero, 1999].
However, the larger fibers have some drawbacks as well. Due to the bulk of material in
the 0.010" diameter fibers, the chances of defects in the fibers is increased. These defects may
lead to fiber cracking.
To compound the problem of fiber cracking, the higher attainable line
fractions limit the amount of epoxy matrix around the fibers. The load transfer path around
cracks is therefore compromised, and cracks may be able to propagate from one fiber to another,
causing macroscopic damage to the composite [Jones, 1999]. Studies have shown that the 0.005"
diameter fiber AFCs retain their mechanical properties to much higher strain levels than the
0.010" fibers, indicating that damage propagation from one fiber to another may be a problem
in 10-mil AFCs [Wickramasinghe, 2000].
2.2.2
Blade Scaling
All work on integral actuation has been performed on scaled down model helicopter blades.
Two scaling methods have been used in designing integral blades:
Froude scaling and Mach
scaling. Froude scaled blades have realistic strains and blade accelerations, but lower dynamic
stress levels than full scale blades [Bielawa, 1992].
Froude scaling has been shown to be best
suited for aeroelastic stability studies [Friedmann, 1998].
A Mach scaled blade has the same
tip speed as a full scale blade. Mach scaling ensures that compressibility effects at the leading
edge of the blade are realistic and that stresses experienced by a model scale blade are equal to
the stresses experienced by a full size blade.
[Bielawa, 1992].
Mach scaling has been shown to
be best suited for vibration control studies of rotor blades [Friedmann, 1998].
2.2.3
MIT/NASA Langley Integral Blade
MIT and NASA built an integrally actuated blade in 1999, known as the Active Twist Rotor
(ATR). This blade was based on a NACA-0012 airfoil, with a chord length of 4.24" and a rotor
radius of 55". The blade was Froude scaled [Cesnik, 1999].
in the NASA Langley Transonic Dynamic Thnnel. [Shin, 1999].
This blade was hover tested
The blade incorporated 24
actuator packs into the spar skin. These packs had 0.005" fibers and screen printed electrodes.
Two plies of actuators were used, one ply oriented at +45' and one ply oriented at -45*.
The
manufacturing procedure for this blade was similar to the manufacturing procedure used in the
32
MIT/Boeing integral blade.
This manufacturing procedure will be described in Section 2.2.4.
The experimentally determined twist rate at 2000Vpp, OVDC was 1.00 /m.
The ATR blade
experienced 5 actuator pack failures during testing, all due to internal pack failure [Shin, 1999].
2.2.4
MIT/Boeing Integral Blade
A 1/6 scale, 60.6" radius, Mach scaled CH-47D integrally actuated rotor blade was built by
MIT and Boeing in 1998.
The blade was built at MIT and hover tested in the MIT hover test
facility.
The CH-47D blade consisted of a D-spar and an aerodynamic fairing.
The spar provides
the structure and the strength of the blade, while the fairing is for aerodynamic purposes only.
The spar layup is seen in Figure ??.
The fairing consisted of a single ply of 450 E-glass with
a few strips of graphite at the trailing edge to add strength and chordwise bending stiffness.
The vertical surface on the back of the spar that forms the interface surface with the fairing is
called the web. See Figure 1-1. The web in the CH-47D blade consists of 3 plies of 45' E-glass
and contributes a significant portion of the blade torsional stiffness [Cesnik,1998].
The CH-47D blade incorporated 42 0.005" fiber AFC packs into the skin of the spar. These
packs were divided into 3 plies, with two plies oriented at
+45* and one ply oriented at -45'.
See Figure 2-1.
During manufacturing of the Boeing/MIT integral blade, electrical connection to 11 of 42
actuator packs was lost.
Several actuators failed during testing.
20 actuator packs were operational.
After initial bench testing,
12 actuators were operational at the end of hover testing.
A full failure history has been published [Rodgers, 1998].
Figure 2-3 shows actuation data for the first generation integral blade.
The blade opera-
tional voltage was limited after several actuator pack failures had occurred to prevent further
damage to the blade.
A half span blade section that was built before the full blade shows the
actuation authority attainable with the design [Rodgers, 1998].
The MIT/Boeing blade showed a 0.80 /m twist rate at 2000Vpp, OVDC, with 69% of actuators working. A Boeing FEM model predicted an actuation level of 1.30 /m for this configuration through the uniform section of the blade, from 0.34R to 0.65R. The model prediction
assumed linear constraint behavior for the material.
33
In other words, the actuation level varied
..........
.
....
. .........
. ..
0.15R
lead-lag pin
~
~packs
~
~
~
~
wire trough
web
active plies
flex circuit
fairing
trailing edge
1.92
4
5.388
IM7 Graphfte @0 dog
E-glass 0 45 dog
AFCO @45 deg
E-glas @ 45 dog
AFCO -45 dog
S-glass @0 dog
AFC 0 45 dog
E-glass @ 45 dog
Figure 2-1: Geometry and layup of the MIT/Boeing integrally actuated rotor blade. The layup
shown represents the layup of the typical active section of the blade. The blade has plies of
+45' AFCs and one ply of -450 AFCs.
Figure 2-2: Integral blade on the MIT hover test stand. AFCs can be seen in the spar skin.
34
.......
... ...................
2.
I
I
I
I
I
2 +
+
1.5
--I
-0 - --
0.5--
00
0
1 B1
ton
da
+v
0
0.5
.
1
1.5
2
2.5
3
3.5
4
Peak to Peak Voltage (kV)
Figure 2-3: Blade section and full blade twist rates. The maximum drive-voltage of the Rodgers
blade was limited to 2000Vpp. A blade section that was built was actuated at 4000Vpp. The
blade data was taken with 69% of the actuators operational. The section data was taken with
10 of 12 actuators operational [Rodgers, 1998]. The model predictions are based on Boeing
FEM predictions. The "corrected" model results incorporate a correction factor to account for
high field non-linearities in the AFCs [Weems, 1998].
linearly with the constraint level of the material.
this assumption may be inaccurate.
However, it was shown by Rodgers that
Individual actuator pack testing had shown that a 0.70
correction factor should be applied to the twist rate predictions based on free actuation levels.
This brought the predicted twist level to 0.910 /m,
only 14% over the measure twist rate at
2000Vpp [Weems 1998].
A transfer function of the Rodgers blade in hover is shown in Figure 2-4. This transfer is
from input voltage to coefficient of thrust, and gives an indication of the actuation authority
through the frequency spectrum [Rodgers, 1998].
While the Rodgers blade showed that the concept of embedding AFCs in a rotor blade was
feasible, the blade experienced numerous actuator failures and therefore fell short of expected
actuation authority.
A total of 34 of the 48 actuator packs had failed or had lost electrical
contact by the end of testing.
11 packs became disconnected during the blade manufacture.
35
10-2
10'
10
4 deg
4 deg
106
0
-500
-1000
-1500
0
100
50
150
Frequency (Hz)
Figure 2-4: Transfer function from input voltage to coefficient of thrust of Rodgers blade in
hover [Rodger, 1998].
36
Another 8 packs were lost due to connection loss or due to failures in the power bus during
testing. The remaining 15 packs failed internally due to short circuits[Rodgers, 1998].
Figure F-1 shows the failure map of the Rodgers blade.
that occurred during manufacturing and operation.
This figure includes all failures
The damage map shows numerous core
voids were observed in the core of the blade when the blade was cut into sections. These core
voids correlated with the location of internal strain gages in the blade indicating that one of
the chemicals used in the blade instrumentation was outgassing during the cure.
The work
that was done to correct the problem of core voids will not be discussed in this thesis, but will
presented in Appendix F.
The loss of the actuator packs in the Rodgers blade can be attributed to immaturity of the
material system and to the difficulties encountered when manufacturing the blade.
The main
source of manufacturing problems were the electrical connections to the actuator packs.
The
connections were made using a solder joint that proved difficult to make because of its location
in the blade and the types of solder used.
Power is supplied to the actuators through a flexible circuit. The flexible circuit consists
of multiple copper lines sandwiched between two layers of Kapton. An acrylic matrix acts as
an insulator between conductive lines. Each copper line is designed to carry up to 0.5 amps at
4000V.
Figure 2-5 shows the flexible circuit design used in the Rodgers blade. The actuator packs
used in the blade had copper electrodes. Before the spar cure, a copper pad was mounted to
the electrode tab to reinforce the tab and reduce the risk of breakage during the cure. This
reinforcement pad was mounted to the electrode tab with a regular lead/tin solder, with a
melting point of 450'F.
The electrode flaps were bent onto the web of the blade during the
spar cure. They were cured in place, and Guaranteed Non-Porous Teflon 2 (GNPT) tape was
placed over the electrode pads to ensure that the conductive surface is left exposed.
Following the spar cure, the flexible circuit was placed against the web and the exposed
copper electrode flaps. The connections were made, supplying heat to the hidden solder connection through an auxiliary copper strip which is trimmed off following the connection. A
low temperature (290'F) indium solder was used to make the connection from the electrode
2
GNPT, American Durafilm Co., Holliston, MA.
37
.......
......................
Auxilliary Copper Strips
For Heat Conduction
Electrode Tabs
Electrode Flap
N
Flexible Circuit
Figure 2-5: Connections in the Rodgers blade are made along the web. The electrode tabs
are bent onto the web, the flexible circuit is placed against the web. Heat is conducted into
the connections through an auxiliary copper strip. The connection is hidden, and can not be
inspected.
X 7 T VX' 2 X'3
"
core void
BS9.093
--
section cut
o
strain gage
>
delamination (inner)
<
delamination (outer)
*
pack electrical failure
>
electrode d
5~i±~iLA(iJ
X 4~
I
-4 M - M A M--7--M --)V Z)V i xe,
UPPER
I
17
21
25
29
33
37
41
45
49
53
57 60.619
5
8
8""
m
Co re
4
BS).H'I0
1
7
21
5
29
3
7 141
45
49 13
LOWER
Figure 2-6: Damage map of the Rodgers blade. [Rodgers, 1998]
38
5
6C.619
tab to the flexible circuit.
This solder was used to prevent melting of the solder joint between
the reinforcing pad and the electrode tab.
However, this solder also had a lower strength
than the regular solder, making the connections prone to failure. Following the flexible circuit
attachment, the fairing was assembled and cured, completing the integral blade [Rodgers, 1998].
The manufacturing procedure described above may have lead to broken connections or short
circuits between connections or power lines.
First, the low temperature solder joint between the electrode tab and the flexible circuit
may have broken during manufacturing and operation because of the lower strength. Also, the
low temperature solder may have been melted by localized overheating of the blade molds. The
blade is cured at 250'F, but because of inadequate thermal control, it is believed that significant
localized overheating may have occurred in the blade mold, with spots reaching temperatures
as high as 300'F. The connections made with the low temperature solder may therefore have
melted and come apart.
Second, the electrode flaps were bent onto the web, making connection to the flexible circuit
difficult. The flexible circuit was brought against the web, hiding the connection area.
Heat
conduction through the auxiliary copper strip into the connection area may have been inadequate. This scenario also prevented inspection of the solder joint. As soon as some part of the
solder joint was made, it became impossible to inspect the connection.
Proper solder melting
and flow could not be verified.
The sharp bend in the electrode tab may have created a stress concentration in the electrode
tab.
This stress concentration may have damaged the electrode tab during the cure or during
operation.
Third, the flexible circuit consisted of six separate layers bonded together. While each layer
of the circuit is very thin and flexible, the bonding created a thick laminate that was very
stiff.
The flexible circuit structure could therefore exert large forces on the electrode tabs
and connections during the cure. Mold pressure or thermal expansion of the circuit may have
caused the circuit to move relative to the electrode tabs, damaging the relatively weak indium
solder connections.
Further, since the web width of the model blade was only 0.6", all the connections were very
closely spaced.
Six flexible circuit layers were used, each carrying 14 lines. 0.1" square solder
39
pads had to be accommodated in the flexible circuit, requiring very close spacing of the lines in
each layer. The spacing between the lines was only 0.015".
Small manufacturing defects, such
as voids or line irregularities, could therefore cause an arcing path between lines in the circuit,
leading to failure.
The motivation for this thesis lies in the problems discussed above. Two areas must be
addressed: determination and quantization of the non-linearities of the actuator material, and
blade design and manufacturing.
The non-linearities in the actuation behavior of the AFCs
must be known in order to improve the modeling of the blades.
that a 30% reduction factor should be applied to predictions.
The Rodgers blade assumed
However, further studies must
be conducted to better understand the non-linear behavior of the AFCs.
Knowledge of the
variation with drive voltage of the behavior of the AFCs under constraint will allow for better
predictions of blade performance throughout the operating range.
This investigation will be
conducted in Chapter 3.
The design and manufacturing of the active system of the blade will be studied in Chapter 4.
Of great concern are the actuators that failed during the manufacturing procedure and during
operation of the blade. Eighteen of the 29 actuators that were lost in the Rodgers blade were
lost due to electrical disconnection or due to short circuits in the power bus.
losses should be eliminated or, at least, minimized.
Obviously, such
Since the AFCs are built by Continuum
Control Corporation, internal failures to the AFCs will not be addressed.
Therefore, the
connections to the AFCs and the power bus design are the main focus of the improvement
to the design and manufacturing procedure.
Alternatives to the hidden solder joint must be
found. A connection between the power bus and the AFCs that can be made more reliably and
that can be inspected after it is made must be developed to prevent the problems encountered
with the Rodgers blade connections. Sound, strong connections to the AFCs will reduce the
risk of connection failures, both during manufacturing of the blade and during operation.
40
Chapter 3
Active Fiber Composite Testing
3.1
Motivation
Two aspects of Active Fiber Composite behavior will be investigated in this chapter: laminated
actuation behavior and electrical behavior.
Laminated actuation behavior will be studied to
improve the accuracy of predictions of blade actuation level. Electrical behavior will be studied
to determine the variation of electrical properties throughout the operating range, and to help
predict the electrical current requirements for testing the blades.
Experience has shown that when AFCs are constrained with passive material, the resulting
actuation level can differ greatly from the level predicted by a model assuming linear elastic
constraint of the laminating material and constant induced stress of the active material. These
blade models base the expansion coefficients of the active material on free strain data. More
detail on the basic modeling concept is presented in Section 3.4.
High field non-linearities
of the AFCs are believed to play a role in the discrepancy between predicted and measured
actuation levels.
Correction factors have been used in modeling of test articles to attempt to
account for this discrepancy. The constrained behavior of AFCs must be studied to determine
experimentally the correction factors needed to correct for the high field non-linear behavior of
the AFCs.
The electrical behavior of the AFCs must be studied to quantify and explain two phenomena
that were observed in the AFCs: higher than expected current draw, and self heating at high
drive levels.
The drive voltage and frequency dependence of AFC capacitance and power loss
41
will be determined. These properties will help explain the causes of the AFC heating and high
current draw.
This study is important to the wind tunnel study of the integrally actuated
blades. The level of current draw of the blades must be known to ensure that enough current
is available to run the blades up to desired voltage and frequency levels. Heating of the blades
may cause degradation of the passive, structural composite materials surrounding the AFCs,
posing safety of flight concerns [Weems, 1998].
Therefore, understanding the magnitude of
the heating of the AFCs and the drive levels at which it becomes a concern is important to
protecting the structural integrity of the blades.
3.2
Approach
Lamination testing is performed on two types of AFC packs:
coupons and 10-mil fiber blade packs.
5-mil fiber characterization
Generally, lamination testing involves testing the
actuation of the AFC before and after lamination between plies of E-glass prepreg'.
The
pre-lamination and post-lamination actuation values can be compared to determine the actuation reduction. Linear model predictions can be compared to the reduction in actuation due to
lamination.
From this comparison, appropriate reduction factors to be applied to the models
can be determined.
The characterization coupons, described in Section 3.5.1, are tested to understand the effect
of constraining layers on actuation. E-glass laminating plies are used to constrain the actuation
of the AFCs. The coupons are laminated at different pressures to determine whether processing
could affect the laminated behavior of the AFC packs.
Ten-mil fiber blade packs are studied to determine the behavior of the AFCs embedded in
the blade.
E-glass laminating material is used for these packs as well.
Two plies of E-glass
are bonded to each surface of the AFC to simulate the constraint level expected in the blade.
The correction factors that should be used in the blade models are determined by comparing
linear model predictions to the experimental results.
The electrical behavior of 10-mil packs is studied.
Capacitance and loss factors are deter-
mined by fitting an equivalent resistor-capacitor circuit model to the current draw data of the
'Hexcel E120-155 prepreg cloth, Pleasanton, CA.
42
E__
EYY
G__
Thickness
Average
Free Actua-
tion
0.005" AFC
29.5 GPa
11.6 GPa
4.0 Gpa
0.010" AFC
35.5 GPa
14.5 GPa
N/A
E-Glass
21.2 GPa
21.2 GPa
3.9 GPa
0.0085"
0.0135"
0.0045"
893p- (18 samples)
1484pe (4 samples)
I
I
I
Table 3.1: Nominal low stress AFC and laminating material mechanical properties [Wickramasinghe, 2000]
laminated blade AFCs.
This model allows the lossy and non-lossy components of the current
draw to be separated, allowing determination of capacitance and dielectric loss factor of the
AFCs.
3.3
AFC Properties Review
This chapter will present only results of lamination, current draw, and heating testing.
Me-
chanical properties have been determined by Wickramasinghe, and are presented in Table 3.1.
Table 3.1 shows the nominal, low stress properties of the AFCs and the laminating material.
These stiffnesses may change after being exposed to high strains due to damage to fibers in
the composites.
In addition to mechanical properties, the fatigue properties of the material
have also been studied.
Both electrical and mechanical fatigue have been investigated.
The
AFCs have been run up to 20,000,000 cycles at 3000Vpp without electrical failure, and up to
10,000,000 cycles at 1000±950pLe without mechanical failure or significant actuation reduction.
The actuation under load properties of the actuator material have also been tested and show
that the material retains its actuation capability beyond the range of strains expected during
testing of the blade [Wickramasinghe, 2000].
3.4
Modeling Concept
A simple actuator constraint model is used to illustrate some of the concepts that will be
introduced in this chapter.
This model is not meant to approximate the blade models in any
way, but is meant only to introduce and discuss the concepts that will be presented in Sections
43
Electric field applied
c
Actuator
9
Figure 3-1: A simple constraint model is used to illustrate the linear constrain modeling technique and the reduction factors necessary to correct for non-linear behavior of the AFCs.
Measured Actuation
Constraint
Stiffness, EC
Predicted Actuation
Predicted actuator behavic
Assumed Actuator
Stiffness, Ea
reduced
Reduced actuator
behavior
free
Actual actuator
behavior
Figure 3-2: Actuator and constraint stress-strain plots, showing the derivation of the assumed
The measured free strain is used along with the assumed actuator
(reduced) free strain.
stiffness to predict a constrained actuation level. Since the actuator behavior is not linear, the
predicted actuation level will not match the measured value. The free strain is adjusted to
match the model to the actuation data, keeping actuator stiffness constant.
44
3.5 and 3.6.
The model consists of an actuator layer and a constraint layer.
occurs only in-plane, and no out of plane deformation is allowed.
0
The actuation
The stress in the actuator,
a, in response to an applied electric field, e, is given by,
Ua
(3.1)
= Ea(ea - de),
where Ea is the stiffness of the actuator, Ea is the strain of the actuator, and d is the longitudinal
piezoelectric constant of the actuator. The stress in the constraint layer, ac,
oc = Ecec,
is given by,
(3.2)
where E, is the stiffness of the constrain layer, and e, is the strain of the constraint layer. Setting
e = Ea = e, and balancing the forces on the actuator and constraint layers, the laminate strain,
e, is found as,
e = de
EaAa
EaAa + EcAC'
(3.3)
where Aa is the crossectional area of the actuator and A, is the crossectional area of the
constraint layer.
The piezoelectric constant used in the model, d, is found by measuring the free strain of the
actuator, efree, and defining d as,
d = Efree
(3.4)
The value of d is then used in Eq. 3.3 to find the laminate strain based on material properties,
laminate geometry, and applied electric field.
Unfortunately, the AFCs do not behave linearly, and the constrained actuation, e, tends
to be lower than that predicted by Eq. 3.3, especially at high fields.
Since the blade models
are linear and can not take into account non-linear behavior of the actuators, the piezoelectric
constant of the actuator, d, must be adjusted to account for the non-linearities in the actuator.
The reduced piezoelectric constant is found by determining a reduced free strain, Ereduced, that
will give an actuation prediction that matches the experimentally determined laminate strain.
45
The reduction factor of the free strain is the percentage reduction of the free strain necessary
to make the model fit the data.
See Figure 3-2.
Fortunately, the constrained behavior of AFCs can be determined by studying individual
packs, and does not need to be done on entire blade structures.
By constraining AFCs at the
level of constraint expected in the blades, the free strain reduction factors can be experimentally
determined at various levels of actuation.
These reduction factors can then be applied to the
blade models to allow for realistic predictions of actuation levels.
3.5
Coupon Lamination Testing
3.5.1
Test Samples
The AFC actuators used for testing the effects of cure pressure, temperature, and lamination
are standard material characterization coupons packs.
These coupons are designed to be
compatible with ASTM standards for materials testing [Wickramasinghe, 2000]. The actuators
have an active length of 3.95", and an active width of 0.5".
0.0085".
The thickness of the packs is
The geometry of the testing coupon is shown in Figure 3-3.
The fibers used in the packs are PZT-5A ceramic, 0.005" fibers. The electrodes are screen
printed silver ink electrodes on 0.001" Kapton.
A film adhesive matrix is used [Pizzocchero,
1999].
Due to fiber delivery limitations, the actuator vendor, Continuum Control Corporation 2,
delivered the AFC coupons to MIT in 3 separate shipments.
Each AFC shipment (batch)
consisted of AFCs manufactured with fibers from a separate fiber shipment from the fiber
vendor, CeraNova 3 . The first two batches consisted of 6 packs each. The third batch included
8 packs.
3.5.2
Test Procedure
The testing procedure consists of three main steps :
testing of the free AFC, processing, and
testing of the laminated AFC. The free AFC testing includes capacitance and actuation testing.
2
3
Continuum Control Corporation, Billerica, MA., 978-670-4910.
CeraNova Corporation, Franklin, MA.
46
6.50 in [1 65.0 mm]
86
[m]
0.50 in [12.5
3.95 in 1100.0 mm]
4.50 in [115.0 mm]
-.
5in[.
m
5.50 in [1 40.0 mm]
Figure 3-3: AFC standard test coupon. The active area of the coupon is 0.50" wide, and
3.95" long. The electrodes are 0.0075" wide, and are spaced 0.045" apart, on center. The
electrode pattern is screen printed silver ink, and the fibers are 0.005" diameter PZT-5A fibers
[Wickramasinghe, 2000].
4
Capacitance of the AFC is measured using a Hewlett-Packard 4194A Impedance/Gain-Phase
Analyzer, at 400Hz excitation frequency. Actuation data is collected for each pack.
testing is performed using a laser interferometer system.
This
A full description of this testing rig
has been published by Wickramasinghe [Wickramasinghe, 2000b].
Actuation data is gathered
at 2000Vpp, OVDC, and 3000Vpp, OVDC, at a frequency of 1Hz.
Three different processes are performed on the AFC coupon packs. Four packs are laminated
at 50psi, 250*F.
cycle only.
Six packs are laminated at 500psi, 250*F. Four packs are put through a heat
These four packs are heated in an oven to 250*F, for 2 hours.
packs are put through no processing.
The remaining 4
Table 3.2 shows the allocation of the packs from the
three different batches.
The laminating pressure of 50psi is the standard cure pressure for the E-glass used for
4
4194A Impedance/Gain-Phase Analyzer, Hewlett Packard, Palo Alto, CA.
Batch
1
2
3
Total
#
Only
Control (#
Heat
of packs)
1
1
2
4
(# of packs)
1
1
2
4
50psi lamination
500 psi lamination
(# of packs)
1
1
2
4
(# of packs)
2
2
2
6
Table 3.2: Test matrix for lamination testing of coupon AFCs. Six packs are put through a
500psi lamination process. Four packs are put through a 50psi (normal) lamination process.
Four packs are exposed to heat only, and four are set aside as control specimens.
47
lamination.
500psi was chosen as an alternate laminating pressure to investigate the effect on
the AFCs of high cure pressures. In the blade cure, much higher pressure than 50psi may be
present. Eight C-clamps rated to 30,000lbs each are used to close the mold. The possible cure
pressure in the blade mold is therefore potentially very high.
The maximum possible value is
not easily calculated because of the complex geometry of the mold and the lack of knowledge of
contact points between the two mold halves. To ensure that the coupon AFCs are not critically
damaged during the lamination process, the AFCs are laminated at a maximum of 500psi. This
is also near the upper operating force for the hotpress that was used. The lamination procedure
is described in Section 3.5.3.
The heat-only cycle was performed on some packs to quantify the effects of heat on the
AFC actuation. It is important to be able to separate temperature induced effects from the
cure pressure induced effects.
The AFCs are held at 250'F for 2 hours.
The Curie tem-
perature, the temperature at which the ceramic loses piezoelectric properties, of PZT-5A is
690'F [Berlincourt, 2000]. Since 250'F is far below this temperature, significant damage to the
piezoelectric properties is not expected, but some changes may occur due to the long duration
of the elevated temperature.
Following processing, actuation and capacitance testing are again performed, following the
procedures described above.
3.5.3
Lamination Procedure
Following initial actuation testing, the AFCs are cleaned with acetone and methanol to remove
any oil or other residues from the surface of the actuator.
This is done to ensure a sound
bond between the laminating material and the actuator surface. Strips of 0.0045" E-glass cloth
prepreg are cut to 0.5" by 5.5", at a 0/900 orientation to the fibers.
The AFCs are then
laminated with one strip of E-glass on each surface, taking care to cover the active region
completely but to not have any E-glass extending over the edges of the AFC coupons. The
AFCs are then placed between two strips of GNPT to protect them from contamination and to
prevent the E-glass from bonding to the cure plates.
Two AFCs are laminated simultaneously between two steel plates. This ensures that the
two AFCs are exposed to the same temperature and pressure, allowing for direct comparisons
48
. . ... ..............
....
...
....
1/16" Rubber
E-glass/A
E-glass
GNPT
Figure 3-4: Setup for coupon lamination in the hotpress. The E-glass/AFC/E-glass laminates
are cured between two steel plates. GNPT is used to prevent the E-glass from bonding to the
cure plate, and a 1/16" piece of rubber is placed on top of the laminated to ensure uniform
pressure distribution through the laminate.
of post-lamination performance and characteristics. A 1/16" thick, 0.5" wide strip of rubber is
also placed on top of each AFC to distribute the load evenly. See Figure 3-4. An aluminum plate
and a piece of soft sponge rubber is placed between the top steel plate and the upper surface of
the hotpress. The sponge rubber is used to ensure that any misalignment between the hotpress
plates does not result in an uneven pressure distribution between the two laminates or within
each laminate. This lamination setup simulates the environment that the AFCs experience in
the blade mold: blade AFCs are laminated between a hard aluminum mold and a relatively
soft spar core.
The hotpress is closed with a force appropriate for generating the desired cure pressure,
50psi or 500psi, and is heated to 250*F. Once the hotpress has reached the cure temperature,
the laminates are cured for 2 hours. The standard cure cycle for the E-glass is 90 minutes at
250 0 F, but to ensure that the heat has transferred properly into the laminates, the cure cycle is
49
extended by 30 minutes. Following the cure, the pressure is released and the heat is turned off.
The laminates and cure plates are allowed to cool in air. This simulates the cooldown period
associated with the blade cure.
After the laminates are removed from the hotpress and the cure plates, excess epoxy flash
is trimmed with a utility knife. Capacitance, resistance, and actuation measurements are then
taken in the manner described above.
3.5.4
Results
It has been shown through a rules-of-mixture model based on the cross-sectional properties of
the AFC that the expected reduction in actuation as a result of lamination with two 0/900
E-glass plies is 50%. This model uses the properties of individual components of the laminate,
including the ceramic fibers, the matrix, the electrode substrate, and the E-glass, to determine
an expected level of reduction in actuation [Wickramasinghe, 2000c].
Table 3.3 shows the reduction in actuation due to heating only. The control specimens show
some reduction as well.
5%.
The experimental error of the actuation testing rig is approximately
The difference between the control and heat-only AFCs is within the experimental error
of the rig, and is therefore negligible.
Table 3.4 shows a summary of the reduction in actuation due to lamination at 50psi and
500psi.
The packs laminated at 500psi show slightly greater actuation reduction percentages.
However, this difference is very small compared to the batch to batch variation in actuation
reduction. Batch 1 performed nearly as expected, whereas Batch 2 and Batch 3 showed greater
than expected reduction in actuation.
The entire data set for these coupons is presented in Appendix A.
This data includes
capacitance measurements and actuation data.
3.5.5
Summary
The main conclusion that can be drawn from this study is that the batch to batch variation in
actuation reduction as a result of lamination dominates any variations caused by the processing
of the AFCs. The coupons that were put though a heating cycle only showed a few percentage
points greater reduction in actuation than the control specimens.
50
The difference between the
Control
Batch 1
Batch 2
Batch 3
Average
2000Vpp
-1% (1)
N/A
-6% (2)
-4%
I Heat Only
3000Vpp
-2% (1)
N/A
-4% (2)
-3%
2000Vpp
-7% (1)
-6% (1)
-5% (2)
-6%
3000Vpp
-5% (1)
-5% (1)
-3% (2)
-4%
Table 3.3: Variation percentage of the actuation level of control coupons and heat-only coupons.
There is only a small difference between the post-processing actuation reduction of the heatonly coupons and the control coupons. This difference is most likely within the margin of error
of the testing rig, and the difference is therefore negligible. The number in parentheses indicates
how many samples the data is based on. The control specimen in Batch 2 was partially depoled
during initial actuation testing. The actuation reduction includes the effect of depoling, and
the data is therefore not shown.
50psi lamination
2000Vpp 3000Vpp
Batch 1
Batch 2
Batch 3
I Average
-50% (1)
-73% (1)
-62% (2)
-62%
-52% (1)
-74% (1)
-64% (2)
-63%
500psi lamination
2000Vpp
-56% (2)
-74% (2)
-68% (2)
-66%
3000Vpp
-57% (2)
-76% (2)
-69% (2)
-67%
[ Average
I 2000Vpp
-54%
-74%
-65%
I 3000Vpp
-55%
-75%
-67%
Table 3.4: Summary of actuation reduction percentages of coupons laminated with different
pressures. The packs laminated at 500psi showed a slightly greater reduction in actuation
than the packs laminated at 50psi. However this difference is much smaller than the batch
to batch variation in actuation reduction. The expected reduction percentage is 50 percent.
Batch 1 performed almost as expected, whereas batches 2 and 3 showed greater than expected
reductions. The number in parentheses indicates how many samples the data is based on.
51
average actuation reduction of packs laminated at 50psi and 500psi is 4%.
The free strain measurements appear to not be sufficient to characterize the behavior of
the AFCs under constraint.
Correction factors must be included in linear models to correct
for these variations, as shown in Section 3.4.
Since the batch to batch difference in actuation
reduction is high, lamination testing should be performed on samples from each batch of AFCs
that is received. This lamination testing will allow for determination of correction factors to be
applied to models of the structures incorporating AFCs from each batch.
3.6
3.6.1
Blade Pack Lamination Testing
Motivation
Section 3.5 showed that there was a great difference in the constrained actuation behavior of
Active Fiber Composites from different batches of packs.
Two batches showed much greater
than expected reduction of actuation due to lamination, while one batch performed almost as
expected.
To accurately predict the actuation level of the integral blade structure, testing
similar to that performed in Section 3.5 must be performed on the blade packs.
3.6.2
Test Procedure
The packs for the Active Material Rotor (AMR) blade are delivered in several batches of 30
packs each.
These packs are 10-mil fiber packs, with the properties described in Table 3.1.
The geometry of the blade packs is shown in Chapter 5 (Figure 5-3).
From each of two batches, batch 1 and batch 3, a pack with high actuation (relative to the
average actuation) and one with low actuation are selected to be lamination tested. Actuators
with both high and low actuation levels are tested to determine if the discrepancies seen in
the coupon testing described in the previous section is a function only of the packs being
manufactured from different batches of fibers, or whether differences in free actuation level also
affect the laminated behavior. Table 3.5 describes the naming convention of the four selected
packs.
Each AFC blade pack is actuation tested in the direction of the fibers, in the middle of the
pack. The location of actuation testing is shown in Figure 3-5.
52
Actuation versus voltage data
Figure 3-5: Actuation testing of the AFCs is performed in the middle of the pack, in the
direction of the fibers.
is gathered for each pack.
Testing consists of gathering actuation data for each pack between
500Vpp and 3000Vpp, at 500Vpp intervals. The actuation frequency is 1Hz.
Lamination is performed in the hotpress in the same manner as the coupon packs.
See
Section 3.5.3. The cure cycle for the E-glass laminating material is 50psi, 250*F for 90 minutes.
The hotpress takes approximately 30 minutes to reach the cure temperature, so the cure cycle
is extended to 120 minutes.
Both packs from each batch are laminated simultaneously, side by
side, to ensure consistency in the cure cycle each pack is put through. Therefore, packs 1L and
1H are laminated simultaneously, and 3L and 3H are laminated simultaneously.
Each pack is
laminated with one ply of E-glass on each surface. The E-glass is oriented at a 0/900 angle to
the direction as the fibers in the packs.
Following lamination, actuation testing is performed again. The AFCs are run at the same
Pack Name
Pack Description
1L
1H
3L
3H
Average
Low free actuation pack from Batch 1
High free actuation pack from Batch 1
Low free actuation pack from Batch 3
High free actuation pack from Batch 3
I
Actuation
Initial
Level at 3000Vpp
1231pte
1803pW
1259pie
1641pe
1484pe
Table 3.5: Description of blade packs that are lamination tested.
53
Pack
Actuation
(Unlami-
Actuation (1-ply lamination)
Actuation (2-ply lamination)
2KV (pe)
3KV (pE)
Model
2KV (pE)
3KV (IE)
Model
509
411
446
506
468
880
704
763
864
802
36%
36%
36%
36%
36%
381
314
329
346
343
647
528
548
574
574
53%
53%
53%
53%
53%
nated)
1H
IL
3L
3H
Average
2KV
3KV
(PE)
(PE)
977
687
706
903
818
1803
1231
1259
1641
1484
(48%)
(40%)
(37%)
(44%)
(43%)
(51%)
(43%)
(39%)
(47%)
(46%)
(61%)
(54%)
(53%)
(62%)
(58%)
(64%)
(57%)
(56%)
(65%)
(61%)
Table 3.6: Summary of results of lamination testing of blade packs. Percentages indicate
reduction percentatges. Although packs 1H and 3H had higher initial actuation, these two
packs show a greater reduction in actuation due to lamination. . The model result is based on
a rules-of-mixture linear elastic model [Wickramasinghe, 2000].
voltage levels as were used for free actuation testing.
Each pack is then laminated with an
additional ply of E-glass to further constrain the AFC. Lamination with a total of 4 plies of
E-glass corresponds to the expected constraint level that the AFCs will experience in the blade.
Actuation data is again gathered and the percentage reduction in actuation at the various
voltages is calculated.
3.6.3
Results
The reduction percentage in actuation due to constraint (lamination) is of importance. This
data indicates whether it is necessary to use correction factors when modeling embedded AFCs
with linear models.
The expected actuation reduction percentage due to lamination with 2 plies of E-glass is
36%.
The expected reduction percentage due to lamination with 4 plies of E-glass is 53%.
These predictions were derived using the one dimensional rules of mixture model developed by
Wickramasinghe [Wickramasinghe, 2000].
3.6.4
Analysis
Figure 3-6 shows that the actuation is non-linear with respect to voltage. As the voltage cycle
in increased, the slope of the actuation-voltage line increases.
In the laminated condition,
the actuation is reduced by the percentage shown in the right-hand plots in Figure 3-6.
54
The
1800
Blade Pack 1H, 1Hz actuation
85
1600
60
1400
55
9 1200
50
C 1000
'
-
2 Ply Lamination Prediction
45
S800
PW-
40
~600
1 Pty Lamination Prediction
400
30
200
25 -
0
0
500
1000
1500
2000
2500
00
30
Blade Pack 1H, actuation reduction
20
50
1000
1500
2000
1800
2000
2500
30
Voltage (Vpp)
Voltage (Vpp)
70
Blade Pack 1L, 1Hz actuation
65
1600
1400
55
(01200
50
0 45
C 1000
800
40
600
35
400
30
0
Lan"A
1 Ply Lammaton Prediction
25
200
0
2Ply Lamination Prediction
500
1000
150
2000
2500
30
Blade Pack IL, actuation reduction
500
200
1000
Blade Pack 3L, 1Hz actuation
60
1001400
1200
55
-
-
2 Ply Lamination Prediction
50
1000
00
45
600
35
40
1 Ply Lamination
iP
Prediction
30-
400
200
0
0
30 0
2500
65
1600-
C
2000
70
2000
1800-
1500
Voltage (Vpp)
Voltage (Vpp)
2P1
26
500
1000
1500
2000
2500
-Blade
0
30
20
Voltage (Vpp)
Pack 3L, actuation reduction
600
1000
1600
2000
30 00
2500
Voltage (Vpp)
70
1800
Blade Pack 3H, 1Hz actuation
60 -
1400
55
1200
50
1000
45
S800
40
<600
35f
400
30-
200
0
0
2
2 Ply Laneation Prediction
1
Ply Lamination Prediction
Blade Pack 3H, actuation reduction
25-
500
1000
1500
I
66 -
1600
2000
2500
3000
Voltage (Vpp)
0
50
100
1500
2000
2500
3000
Voltage (Vpp)
Figure 3-6: Blade pack lamination data. The graphs on the left show the actuation reduction
percentage as a function of voltage. The reduction percentage increases with higher voltage.
Also, packs with higher free actuation levels show a greater actuation reduction.
55
30
25
11
Free st ai reducti n perce
t__rri a
for
-----_
20
15
C10
0
0
-5
__
0
500
1000
1500
2000
Voltage (Vpp)
2500
3000
Figure 3-7: Reduction factors necessary to correct the linear models for non-linear actuation
effects of AFCs. High free strain packs need a higher correction percentage than low free strain
packs.
actuation reduction percentage tends to increase as the voltage cycle is increased.
Actuator packs IL and 3L show near expected reduction percentage as a result of lamination.
See Table 3.6. Actuator packs 1H and 3H, which had significantly higher than average actuation
levels,
show a significantly greater reduction percentage as a result of lamination.
Before
lamination, the actuation values all lie within approximately 20% of the average actuation.
After lamination with two plies, the actuation values all lie within approximately 10% of the
average actuation.
It appears that the actuator packs with higher actuation levels begin to
act more like packs with lower actuation levels when constrained.
This indicates that high
free strain does not translate to high laminated strain, probably because of non-linearities, with
voltage, of the induced stress capability of the actuators.
The average actuation level after lamination with 2 plies of e-glass on each side is 574pE at
3000Vpp. The Wickramasinghe model calculates the reduction from free strain to laminated
strain as 53%. This reduction percentage indicates that a free strain of 1225pE should be used
in the linear models to give a better approximation of the twist performance of the blade at
56
3000Vpp.
This is a reduction of 17% from the actual free strain average of 1483pE.
Figure
3-7 shows the reduction factors that must be applied to low free strain packs (average of 1L
and 3L), and high free strain packs (average of 1H and 3H).
High free strain packs must be
corrected by a much larger reduction factor than low free strain packs.
Enough packs have not been tested to yield a statistically significant data set, but it is
clear that an accurate prediction of actuation can not be achieved without incorporating the
non-linear behavior of the AFCs in the form of voltage dependent reduction factors.
This is
especially true for packs with high free strain.
3.6.5
Summary
Two conclusions can be drawn from this study. First, it was shown that laminated packs do not
behave as predicted by linear models assuming linear elastic constraint of the material and linear
induced stress behavior of the AFCs.
at high voltages.
Rather, the packs tend to perform below predictions
Free strain data of the packs can therefore not be used to determine the
expansion coefficients of the active material in the models. Appropriate reduction factors must
be applied to the free strain data of the packs prior to incorporation into the blade models.
Second, historically, high free strain packs have been preferred to low actuation packs for
embedding in structures.
This study, however, has shown that high free strain packs do not
significantly outperform low free strain packs when laminated.
Therefore, it may not be
necessary to discriminate between high and low free strain actuators when determining which
AFCs to embed in a structure.
Obviously, lamination testing should be performed on several
samples from each batch of AFCs to determine the degree of non-linearity of the actuation
behavior, and to determine the reduction factors necessary to correct the linear models.
3.7
3.7.1
Blade Pack Heating and Current Draw Testing
Motivation
Blade pack self heating and current draw are studied. Since excess heating could cause degradation to the passive plies in the blade, understanding the magnitude of the heating of the
blade packs is important to protecting the structural integrity of the blades.
57
The voltage and
frequency dependence of heating in the AFCs must be studied to determine how to minimize
the risk of dangerous overheating of the blades.
Current draw of the packs must be known to allow predictions of blade current draw to
be made.
Determination of the blade current draw will drive the acquisition of high voltage
amplifiers for the wind tunnel test of the blades.
Since the high voltage amplifiers being
considered have current limits, it is important to know the magnitude of the current draw
of the blades so that enough current can be supplied to drive the blade to desired levels.
In
addition, the dielectric properties of the AFCs must be studied to better understand the voltage
dependence of current draw and the self heating experienced by the packs.
3.7.2
Test Samples
Blade packs are used for both the heating study and the current draw study. These packs are of
the same geometry as the packs used for blade pack lamination testing. Again, details on these
packs can be found in Chapter 5, Figure 5-3. Four unlaminated packs and four laminated packs
are tested.
The unlaminated packs are used for heating testing and for current draw testing.
Laminated packs can not be used in the heating study since external temperature measurements
are not representative of the actual AFC temperature: the thermocouple is removed from the
AFC by two plies of E-glass. Laminated packs are used for current draw testing instead. The
laminated packs are the same packs as those described in Section 3.6, and are constrained with
2 plies of E-glass on each surface, approximating the level of constraint experience by AFCs in
the blade.
3.7.3
Test Procedure
The temperature of the bare AFCs is measured with a differential thermocouple 5 . This instrument consists of two thermocouple probes, one of which is held against the sample, and
the other is left in the ambient temperature. The device measures the difference between the
sample temperature and the ambient temperature.
The heating study is performed on three unlaminated packs. Packs U1 and U2 are driven at
3000Vpp, OVDC, 10Hz, 40Hz, and 70Hz, and at 2000Vpp, 70Hz. Lower frequencies at 2000Vpp
5Model 821 Microprocessor Thernomter, Tegam, Inc., Geneva, OH, 440-466-6100.
58
Pack
Description
3000Vpp actuation
Low field capacitance
Internal Pack Name
U1
U2
U3
U4
IL
1H
3L
3H
Unlaminated
Unlaminated
Unlaminated
Unlaminated
Laminated
Laminated
Laminated
Laminated
1240pe*
1485pE*
15281E*
1250pE*
528pe
647pe
548pie
574pE
10.36nF*
9.89nF*
8.66nF*
9.13nF*
7.64nF
8.58nF
7.56nF
7.97nF
Anj-031
Anj-037
Anj-046
Anj-048
Anj-026
Anj-024
Anj-075
Anj-082
* Continuum Control Corporation data
Table 3.7: Heating and current draw testing packs. Actuation values are free actuation for the
unlaminated packs, and constrained actuation for laminated packs. Internal pack names are
given. These are the names given to the packs by Continuum Control Corporation.
are not performed because of the negligible heating levels that occur at these frequencies.
Temperature data is collected initially every 10 seconds, with the sampling intervals increasing
through the testing period.
Testing is done for 10 minutes, with the exception of the 3000Vpp,
70Hz testing of U2, which is extended to 25 minutes to determine the asymptotic behavior
of the heating.
Packs U1 and U4 are tested at 3000Vpp, 45Hz, with 0VDC, 200VDC, and
500VDC offsets.
These packs are run for 5 minutes at each offset level.
After each test run, the AFC and the testing rig are allowed to cool to within 5'F of room
temperature.
Current draw data is collected using the current monitor on the Trek 663A 6 amplifier that
is used to drive the AFCs.
The dependence of AFC current draw on voltage is studied by
running several packs from 500Vpp to 3000Vpp, OVDC, at 45Hz.
Unlaminated packs U2 and
U3 are used for this testing, as are laminated packs 1L and 1H.
The effect of adding a DC offset to the drive voltage is studied on pack U1.
The pack is
driven at 3000Vpp, 45Hz, with OVDC, 200VDC, and 500VDC.
To further study the electrical behavior of the AFCs through the voltage and frequency
range, voltage and current data are fitted to a parallel capacitance-resistance model. This
testing is performed to determine if the increase in current draw at high voltages is due to increased loss (lower apparent resistance) or due to an increase in the dielectric constant (increase
in capacitance) of the material. The data analysis is shown in Section 3.7.6.
6
Model 663A 10KV, 20mA amplifier. Trek Inc., Medina, NY.
59
90
90
Pack U1 (unlaminated)
80
80
70
70
60
E 50
E 50
LL
Pack U2 (unlaminated)
-K
-p,,40Hz
60
-
40
3KVpp, 40Hz
40
E 30
KKVpp,Hz
E 30
(D
20
20
3KVpp, 10Hz
10
00
100
200
300
500
400
2KVpp, 70Hz
10
3KVpp, 10Hz
0
600
500
0
Time (sec)
1.4
1.4
1.2
1.2
A,)
C
LL
1
3KVpp, 70Hz
N
2:
I.
E
3KVpp, 40Hz
3KVpp, 10Hz
3KVpp, 70Hz
-
1
0.8
N
0.6
aI 0.6
3KVpp, 40Hz
0E
0.41
0.4
0.2
0
1500
E
3KVpp,10Hz
0.8
1000
Time (sec)
0
100
0.2
Pack U1 (unlaminated) Normalized byfrequency
200
300
400
500
600
Time (sec)
0
Pack U2 (unlaminated)
Normnlizrd hy frnunrv
0
1000
500
1500
Time (sec)
Figure 3-8: Time histories of heating of U1 and U1. The temperature reaches steady state
after approximatly 15 minutes. The bottom plots show the curves normalized by temperature.
The curves collapse, indicating that the heating is proportional to frequency.
3.7.4
Heating Results and Analysis
Time histories of actuator pack temperatures are shown in Figure 3-8. The data shown is for
unlaminated packs U1 and U2. It can be seen that as the frequency increases, the heating of the
packs increases.
Figure 3-8 also shows the heating data normalized by frequency. The heating
curves essentially converge, indicating a linear relationship with frequency. The actuator packs
appear to approach a steady-state temperature after 10 to 15 minutes. The heating in the packs
can be substantially reduced by lowering the voltage. The heating at 2KVpp, OVDC, 70Hz is
only slightly higher than the heating at 3KVpp, OVDC, 10Hz, and is five times lower than the
60
60
0
46 --
46 -
S40 -
30
OVDC
40 -
35
OVDC
.
-300DC
~3 E-
200VDC
20OVDC
U-26
0
0
U4 3000Vpp, 45Hz
U1 3000Vpp, 45Hz
0
50
100
150
200
260
300
0
50
100
Time (sec)
150
200
250
300
Time (sec)
Figure 3-9: Effect of DC offset on blade pack heating.
by 30% to 40%.
A 500VDC offset reduces the heating
heating at 3KVpp, 0VDC, 70Hz.
The heating can also be reduced by applying a DC offset to the excitation signal. Figure
3-9 shows heating time histories for packs ajn-031 and ajn-048. As the DC offset in the signal
is increased to 500VDC, the heating is reduced by 30% to 40%.
3.7.5
Current Draw Results
Historically, the current draw of the AFC packs has been assumed to approximate the behavior
of an ideal, linear capacitor.
The maximum current draw of an ideal capacitor is given by,
imax
= 27rCAf,
(3.5)
where imax is the maximum current drawn by the capacitor, C is the capacitance, A is the
amplitude of the voltage input, and f is the driving frequency.
The current draw data collected for the unlaminated packs (U2 and U3) and for the laminated packs (1H and 1L) is presented in Figures 3-10 and 3-11.
In unlaminated packs U2
and U3, the current level at 500Vpp is between 1.5 and 2.0 times greater than expected. At
3KVpp, the current level is between 4.5 and 5.0 times greater than expected. In the laminated
packs, the current levels are slightly lower. At 500Vpp, the current level in laminated packs 1H
61
.
0.02
U2 (unlaminated) current draw
45Hz, OVDC
-0015
-.
015
U3 (unlaminated) current draw
.4 5Hz, OVDC
0.01
0.01
Model
0.005
0
500
1000
1500
Voltage (Vpp)
2000
0.005
-
2500
0
3000
Model
- -500
1000
1500
2000
2500
3000
2500
3000
Voltage (Vpp)
Actual/Model
4-
Actual/Model
32-
0
500
1000
1
1500
2000
3000
2500
0
52
500
1000
Figure 3-10: Current draw of unlaminated packs U2 and U3.
than the draw expected based on a linear capacitor model.
2000
Current draw is much larger
A
^^%
0.02
S 0 .015
1500
1 H (laminated) current draw
45Hz, OVDC
0.015
1 L (laminated) current draw
45Hz, OVDC
C
0.01
e
00.005
0.
0
0.01
-
00.005
Model_
500
1000
1500
2000
2500
io"00
C
Model
Soo
1000
Actual/Model
4
3
3
2
2
1000
2500
3000
2500
3000
Actual/Model
4
500
2000
5
5
0
1500
Voltage (Vpp)
Voltage (Vpp)
1500
2000
2500
3000
0
500
1000
1500
2000
Figure 3-11: Current draw of laminated packs 1H and IL. Current draw is larger than expected
based on the linear capacitor model, but smaller than that seen in U2 and U3. The constraint
of the laminating material may be causing a reduction in charge displacement.
62
0.016
U2 (unlaminated)
Current vs. Frequency, OVDC
0.014
-
0.012 0.01
S0.008
2 0.006
0.004 0.002
10
20
30
40
50
60
70
Frequency (Hz)
80
90
100
Figure 3-12: Current vs. frequency data for unlaminated pack U2. Peak current draw varies
linearly with frequency, allowing reasonable extrapolations to be made from given data to higher
frequencies.
and 1L is approximately 1.5 times higher than predicted, and at 3KVpp, the current level is
between 2.5 and 3.5 times greater that expected. The behavior seen in the AFCs at high fields
is consistent with published data that indicates that ferroelectric materials are very non-linear
[Braithwaite 1990; Jaffe, 1971; Berlincourt, 2000].
The current draw increases linearly with frequency, as seen Figure 3-12. Therefore, although
most data is taken at 45Hz, reasonable extrapolations can be made to determine current draws
at higher or lower frequencies.
The addition of a DC offset to the drive signal can decrease the current draw of the AFCs.
Figure 3-13 shows the magnitude of the decrease in maximum current level for an AFC being
run with 0, 200, and 500VDC offset. The maximum current draw can be decreased by 20% by
applying a DC offset of 500VDC to the drive signal.
3.7.6
Power Loss
It is known that as driving fields increase, the dielectric constant of the piezoelectric ceramic
increases, and the material becomes more lossy [Jaffe, 1971]. The losses are associated primarily
with hysteresis of the dielectric displacement-electric field relationship. The hysteresis is caused
by electrical domain reorientation within the ceramic [Berlincourt, 2000b].
63
Soft ceramics, such
0.02
0.01520OVDC
0.01 -
SOOVDC
0.005OVD)
0.005 -
0.01
AJN031, 3KVpp, 45Hz
0.015
2000
1500
1000
500
0
500
1000
1500
2000
2500
Voltage (V)
Figure 3-13: Current vs. voltage for various DC offsets on unlaminated pack U1. The current
levels can be reduced by 20% by adding a 500VDC offset to the excitation voltage.
as PZT-5A, of which the AFC fibers are made, are especially prone to dielectric losses and these
losses can cause significant self heating [Berlincourt, 2000].
It is important to characterize the dielectric loss and dielectric constant of the Active Fiber
Composites to determine the cause of the higher current draw observed in the blade packs.
Higher than expected loss and dielectric constant, and therefore capacitance, can both contribute to the higher current draw.
To find the power loss of the AFCs,
the input current can be multiplied by the input
voltage and averaged over an integral number of cycles.
This method, however, does not
provide information about the non-dissipative characteristics of the material.
To gain insight
into the dielectric change as a function of voltage and frequency, as well as the magnitude of
the dissipative characteristics of the material, the current and voltage data is analyzed using
an equivalent electrical circuit model.
This model consists of a capacitor in parallel with a
dissipative element.
The portion of the current draw that is out of phase from the input voltage is non-dissipative.
The portion that is in-phase with the input voltage is dissipative.
By modeling the AFCs using
a capacitor-resistor model in parallel, the dissipative and non-dissipative electrical behaviors
can be separated.
The capacitor captures the changes in capacitance of the material, while
the dissipative element (or resistor) captures the energy dissipation of the material. From the
64
07
Figure 3-14: Equivalent circuit model for the AFCs. The capacitor captures the non-dissipative
electrical behavior, while the resistor captures the the dissipative behavior.
dissipative properties, presented as equivalent resistances, the power loss of the material can be
determined.
It should be noted that it is assumed that mechanical losses are negligible, and
that the power loss in the AFCs is dominated by the electrical behavior of the piezoelectric
fibers.
The current running through RC circuit shown in Figure 3-14 in response to a voltage
excitation,
V = Asinwt
(3.6)
1
i = A|Clwcoswt + -- A sin wt,
R
(3.7)
is,
where v is the excitation voltage, A is amplitude of the voltage signal, W is the frequency in
radians per second, t is time, i is current, JCJ is capacitance, and R is resistance.
The Fourier coefficients of the current draw of the AFCs at the driving frequency are found.
The data is manipulated to find the magnitude of the current signal, and the phase of the
current signal with respect to the voltage signal. The current is written as,
i = B sin(wt + #),
65
(3.8)
4
00
a <10
3 10
1H (laminated)
1 H (laminated)
1H (laminated)
6
0015
S0.01
0
0
t(v)
C0
(3g
(e0(
-2
-1
-,000-
behavior
-Model
-2
AF behavior
Figure 3-15: Model fit to data. The fit captures the current curve characteristics well through
2000Vpp. As voltage approaches depoling voltage, the material begins to behave non-linearly,
and the details of the voltage-current curve can not be captured by a R-C model.
where B is the current amplitude and
voltage.
#
is the phase difference between the current and the
ICI and R are found by trigonometric identities relating the form of Eq. 3.8 to the
form of Eq. 3.7,
B =
=
#
(AICjw) 2+ (-) 2,
(3.9)
(3.10)
tan-1(RClw).
The values for the resistance and the capacitance are found as,
ICI
R =
B
AB
tanq5
tan
A
-
1 + tan2 #.
Aw V/1+ tan24
B
'2(3.11)
(3.12)
The current versus voltage data for several drive levels of a laminated AFC are shown in
Figure 3-15.
The fit of the model to the data is shown. The fit captures the characteristics of
the current-voltage curve well through 2000Vpp. Above 2000Vpp, the AFCs begin to approach
the coercive field and the electrical behavior becomes more non-linear. The linear model begins
66
to miss some of the features of the voltage-current curve at these high driving voltages.
The results of the modeling are shown in Figure 3-16. The equivalent resistance decreases
and the capacitance increases as the voltage increases. The resistance data is an intermediate
step necessary to find the power loss through the dissipative element in the model. However, it
is not a "real" resistance.
The ceramics have very high steady state resistance, and dissipation
through conduction of the material occurs on the timescale of tens of minutes, and will not
have significant effect when running the AFCs at 10 to 50Hz [Harper, 1999].
Dielectric loss is
not a conductive loss, but rather a loss associated with the molecular behavior of the dielectric
material itself.
Figure 3-17 shows the percentage of the total current that runs through the dissipative
path in the model.
As the voltage is increased, this percentage increases significantly, from
approximately 15% to approximately 25%.
The power draw of the equivalent circuit is defined as,
2
_(V(t))
P(t) =
(3.13)
,
where Vrm, is the root mean square value of the input voltage, and Pavg is the average power
output of the equivalent circuit. Z, the impedance of the circuit, is defined as,
Z=
R
.
1+ iwRIC
(3.14)
where w is frequency. Combining Eqs. 3.13 and 3.14,
P(t) =
R$(
1
+ iwRICI),
(3.15)
and taking the time average of the real portion of the power, thereby omitting the non-lossy
reactive power draw, leaves,
V2
Pag =
< Re(P(t)) >
= Vim.
R
(3.16)
The resistance data shown in Figure 3-16 is used to calculate the power dissipation of
the AFCs.
See Figure 3-18.
The power can be seen increasing drastically as the voltage is
67
1H Resistance
2.8
10
2.0
2.4
2.2
30Hz
0
2
a
10Hz
a
10Hz
50Hz
1.0
4
1.6
1.4
30Hz
2
1.2
50Hz
1 H Capacitance
0
500
1500
1000
2000
2500
0
30C
500
0
Voltage (Vpp)
1000
1500
2000
3000
2500
Voltage (Vpp)
x10
010
iL Resistance
10Hz
2.8
10
2.6
2.4
a
2.2
2 1.8
30Hz
1.4
50Hz1
1.2
0
500
1000
1500
.1
10Hz
106
4
30Hz
50Hz
L Capacitance
2000
2500
0
30 00
500
1000
Voltage (Vpp)
1500
2000
2500
300
Voltage (Vpp)
X10
10
2.8
2.0
2.4
3L Resistance
10OHz
10
8
2.2
30Hz
106
0Hz
4
30Hz
50Hz
1.4
2
1.2
50Hz
3L Capacitance
0
cc
3
0 2.8
500
1000
1500
2000
2500
30
0
0
Voltage (Vpp)
C
410
e
Soo
1000
1500
20DO
Voltage (Vpp)
2500
3020
12 0
3H Resistance
10
2.0
24
a
2.2
30Hz
- 10Hz
10Hz
2 1.
C)
50Hz
-
108
2
3H Capacitance
1.2
1 0
500
1000
1500
2000
2500
0L0
3000
Voltage (Vpp)
30Hz
50Hz
500
1000
1500
2000
2500
3000
Voltage (Vpp)
Figure 3-16: Capacitance and resistance for laminated packs 1L, 1H, 3L, and 3H. The high free
actuation packs (1H, 3H) show a greater increase in capacitance than the low free actuation
packs. The resistance data is used to find the power draw.
68
%of total current drawn by resistor inR-C model
30
1H
28
26
26
24
24
50
20
10Hz
-
18
-
16
14
14
12
12
S
1000
1500
2000
2500
30HZ
10
30 00O
Soo0
1000
28
total current drawn
by resiator in R-C model
26
24
24
22
22
-
50Hzz
.
2000
2500
300
current drawn by
resistor in R-C
2500
3000
model
3H
28
26
50Hz
-
.
10Hz
20
18
18
.
16
14
12
12
0
~
500
1000
1500
2000
2500
-
16
10Hzz
14
10
% of total
30
3L
20
1500
Voltage (Vpp)
Voltage (Vpp)
% of
5Hz
10HZ
18
16
0
inR-C model
22
510Hz
20
10
current drawn by rosor
1L
28
22
% of total
30-
10
3000
0
50
30Hz
1000
1500
2000
Voltage (Vpp)
Voltage (Vpp)
Figure 3-17: Percentage of total maximum current that runs through the dissipative path in the
idealized R-C circuit. As the voltage is increased, a greater portion of the current is dissipated.
69
4
1L
1H
3.5
3 .5
3
3
-
2.5
2 .5
2
0
0.
a.
50Hz
:2
50Hz
1.5
30Hz
30Hz
10Hz
0.5
0 .5-
10Hz
0'
0
5O
1000
1500
2000
2500
0
30 00
0
500
1000
4
4
2000
3.5
3. 5
3
3 --
2.5
2. 5
a
0 S2
2 1.
50Hz
50Hz
1.5
5z
15-
1
30Hz
30Hz
0.5
0 .5
10Hz
-1Hz
So0
30( )0
2500
-3L
3L
0
1500
Voltage (Vpp)
Voltage (Vpp)
500
1000
1500
2000
2500
S0
3000
Vokage(Vpp)
500
1000
1500
2000
2500
3000
Voltage (Vpp)
Figure 3-18: Power draw data for laminated packs IH, 1L, 3L, and 3H. The power draw
increases rapidly with voltage.
The high free actuation packs, 1H and 3H, show a much
greater power loss than the low free actuation packs.
70
0.5
1L
1H
0.45-
0.45
0.4-
0.4 I-
30Hz
0.35-
0.35
0.3
C
50Hz
-
0.25.
30Hz
0.3
0.25-
1 O.
50Hz
0.2 [
0.2-
10Hz
0.15-
0.15-.
'
0.1
0
500
'
1000
'
1500
'
2000
:
2500
0.1
3000
0
500
1000
1500
0.5
1
30 DO
0.5
3L
D.45
0.45
0.4
3H
-
0.4 -
0.35
30Hz
.35 '0
C
0.3
0
30Hz
0.25
0.3
50Hz
10H
1.25
50Hz
0.2
0.2
D.15
0.15
0.1
0
2500
Voltage (Vpp)
Votage (Vpp)
C
C
2000
500
1000
1500
2000
2500
3000
0
-
500
1000
150
2000
250
3000
Voltage (Vpp)
Voltage (Vpp)
Figure 3-19: Tan6 data for laminated packs 1H, 1L, 3L, and 3H. High free strain packs, 1H
and 3H, show a greater increase in the dielectric loss factor than low free strain packs.
71
increased.
The drastic increase in heating at 3000Vpp over heating at 2000Vpp in Figure 3-8
can be explained by the large increase in power output at high voltage.
This power output is
approximately 3 times larger at 3000Vpp than at 2000Vpp.
Alternatively, it is common practice to represent the capacitance, and more fundamentally
the dielectric constant of the material, as complex property. Values of the dielectric constant
and of the dielectric loss factor, tan6, are usually presented as part of the fundamental properties
of a piezoelectric ceramic [Berlincourt, 2000; Jaffe, 1971]. Representing the electrical properties
of a ceramic as a complex quantity allows analysis of the power draw of the material without
utilizing an equivalent circuit, but by describing the material in terms of complex properties.
The complex dielectric constant of the material is given by,
T=
where leT
IeT I(1 - i tan 6)
(3.17)
is the magnitude of the dielectric constant, and tan6 is the dielectric loss factor
[Glenn, 2001].
The capacitance, C, of the material is related to E3 by,
C = EA.
t
where t and A are thickness and area quantities.
Therefore,
C = 1CI(1 - i tan 6)
where CI is the magnitude of the capacitance.
Z=
(3.18)
(3.19)
The impedance of the material is,
1
iWC
,
(3.20)
and by Eq. 3.13, the power draw of the material becomes,
P(t) = (V(t)) 2 wICItan 6 + iWCIV2,
(3.21)
where the real term gives the real (dissipative) power draw, and the imaginary term gives the
reactive (non-dissipative) power draw. Taking the time average of the real power draw leaves,
72
Pavg=
<Re(P(t))> =V,2,w|Cjtan6.
This real power draw corresponds to the expression for the dielectrically dissipated heat power
density, PDE, given by Berlincourt and Krueger [Berlincourt, 2000b],
PDE =
we2mse
(3.22)
tan 6,
where erm, is the RMS electric field.
The dielectric loss factor, tan 6, a material property describing the lossiness of the system,
can be found by analyzing the phase difference between the input voltage and the current,
.
The loss in a piezoelectric ceramic is defined as the ratio of the out-of-phase component of the
charge in the ceramic (lossy component) to the in-phase component of the charge (non-lossy
component)
[Jaffe, 1971]. Since 0 is the phase difference between current and voltage, # must
be rotated to correspond to the phase difference between charge in ceramic and voltage. 6 is
related to 0 by,
7r
6 = - - #
2
(3.23)
Figure 3-19 shows tan 6 as a function of drive voltage. The tan6 of the AFCs is relatively
insensitive to frequency, but varies greatly with voltage. The dielectric loss factor increases by
approximately 100% from low field to high field drive levels.
Eq. 3.21 offers insight into the power dissipation as a function of the material properties
0C1,
the capacitance magnitude, and tan6, the dielectric loss factors. It should be clear from Figures
3-16 and 3-19 that both IC and tan6 vary significantly with voltage.
The large increases in
both parameters create the significant amount of heating observed in the AFCs.
Eq. 3.22 should be used as one of the metrics that influences the selection of fiber material.
Readily available information on eT and tan6 should be used to determine the level of heating
that is expected of a given piezoelectric ceramic [Berlincourt, 2000; Berlincourt 2000b].
The
heat power generation of a material should be minimized in order to avoid heat induced damage
to the AFCs and the structure in which they are embedded.
73
Based on published information
0.014-
24
0.012-
0.01
Average maximum current draw for
1 H, 1 L, 3L, 3H, 45Hz excitation
200
mA
1200nA
-
-
120
-
150
___
__
0.008 800m
E0.006
-
--
100 -
T --
0.004
0.0020
T
T-
---
-
500
10
1500
2000
Voltage (Vpp)
2500
3000
500
0
1000
1500
2000
2500
3000
Voltage (Vpp)
Figure 3-20: The plot on the left shows the average current draw of the 4 laminated packs (1H,
1L, 3L, 3H). The plot on the right shows the performance envelope of a 3 blade set, based
on the average current draw data. Running with 3 400mA amplifiers, the highest frequency
attainable at 3000Vpp is approximately 70Hz.
(See Appendix E),
using PZT-4 fibers could significantly reduce the increase in current draw
and loss at high drive levels.
PZT-4 has significantly lower dielectric loss factor and dielectric
constant than PZT-5A. However, it also has a lower piezoelectric constant, which would sacrifice
actuation performance.
3.7.7
Blade Current Requirements and Performance Envelope
Figure 3-20 shows the average current draw at 45Hz, 3KVpp, OVDC of blade packs 1H, 1L, 3L,
and 3H, each laminated with 2 plies of E-glass on each side. Based on these current values, a
voltage and frequency envelope can be found given an available current. The amplifiers that will
be used for blade windtunnel testing are Trek 7 923A amplifiers. These amplifiers can supply a
maximum of 400mA and can run up to
+/- 2KVpp.
The envelopes shown assume a 3 blade
set, drawing the maximum current from the amplifiers.
Heating of the blades should be kept to a minimum. Convective heat transfer during hover
testing should keep the blade cool.
Problems may occur during high voltage bench testing.
Under no circumstances should the blade be run at above 120'F.
7
Model 923A 2KV, 400mA amplifier. Trek Inc., Medina, NY.
74
Above that temperature,
decomposition of the passive material surrounding the AFCs could occur.
3.7.8
Summary
The work presented in this chapter has shown that piezoelectric materials are very non-linear
in their mechanical and electrical behavior. Three important conclusions can be drawn.
First, reduction factors are necessary in the blade models to account for non-linear material
behavior, especially at high fields.
Generally, the actuation predicted by a linear model,
assuming constant actuator stiffness and a piezoelectric constant based on the free strain of the
actuator, is greater than the measured actuation. See Section 3.4. Reduction factors adjust the
free strain that is assumed in calculating the piezoelectric constants used in the models. This
is done to compensate for non-linearities in the electromechanical behavior of the actuators.
These reduction factors must be experimentally determined as they differ from pack to pack
and from batch to batch. Several packs from each batch of AFCs must be lamination tested,
as described in Sections 3.5 and 3.6, to derive appropriate reduction factors to be used in the
models.
For high field (3000Vpp) actuation predictions of the 10-mil fiber AFCs tested in this chapter, the free strains used in modeling the laminated actuation should be reduced by between
10% and 25% from the actual free strains.
Packs with free actuation levels of 1600-1800pe
were shown to require a correction factor of 25% and packs with free actuation of approximately
1250pE required correction factors of about 10%.
Second, the electrical behavior of the AFCs is very non-linear. The AFCs draw significantly
more current than is predicted using a linear capacitor model. This high current draw can be
attributed to a large increase in dielectric constant and dielectric loss factor as a function of
driving field. The current draw of the AFCs can be reduced by the addition of a DC offset to
the drive voltage.
This DC offset ensures that the AFC is not driven near the coercive field.
The large currents associated with depoling of domains in the ceramic are therefore not seen.
Third, blade heating will likely be a problem at high driving voltages. At high driving fields,
the dielectric loss factor of the AFCs increases, leading to higher power dissipation, causing heat
generation in the AFCs.
The AFCs in the blade are embedded in glass and epoxy, both of
which will tend to insulate the AFCs, possibly compounding the heating problem.
75
Again,
the addition of a DC offset to the driving voltage will decrease the current draw of the AFC,
decreasing the heating experienced by the actuator.
The various phenomena described in this chapter will be tested on a blade to determine to
what degree these factors manifest themselves in a realistic blade structure. This work will be
presented in Chapter 5.
76
Chapter 4
Design and Manufacturing of
Internal High Voltage Connections
4.1
Motivation
The manufacturing difficulties in the first generation blade were due to the flexible circuit and
all the connections lying in the web of the blade, away from the surface.
As described in
Chapter 2 all the connections were made against the web. The connections were manufactured
with a " blind" solder joint: heat had to be conducted into the joint through an auxiliary copper
strip.
It was therefore impossible to inspect the quality of the joint after it was made.
The
manufacturing of these electrical connections proved difficult, and several electrical connections
failed during manufacturing and during operation.
These failures may have been caused by
either connections opening up, or by short circuits between the closely spaced connections. The
power bus, including the flexible circuit used to conduct power from outside the blade to the
packs, and the connections to the packs, must be redesigned to improve manufacturing of the
electrical connections, and to make them less susceptible to failure during manufacturing and
testing.
A change in the actuator system also makes it necessary to change the manufacturing
procedure of the power bus.
Screen printed, silver ink electrodes were substituted for the
copper-Kapton electrodes used in the past, as discussed in Section 2.2.1.
This complicates
manufacturing because solder can not be used to make electrical contact with the electrodes.
77
.
...
...........
..
....
...
Flex exits at root
Flex transition
Actuator Packs
Flex transition
Copper Strip
Inboard, upper layer of
flexible circuit
Solder Joint
Spar
Outboard, upper layer of
flexible circuit
Fairing
Figure 4-1: Active materials rotor blade showing the routing of the flexible circuit. Also shown
are the connections from the AFCs to the flexible circuit.
Modifications must be incorporated into the power bus design to accommodate this change. It
is necessary also to take additional steps in the manufacturing process of the actuator packs to
allow for connection to the flexible circuit. This will be discussed in Section 4.3.
4.2
Approach
Ease of manufacturing and robustness of the connections between the actuator packs and the
flexible circuit is the primary concern in the redesign of the power bus.
The power bus
is redesigned to bring the connections to the surface of the blade, rather than having the
connections buried in the web.
The flexible circuit is designed to enter the blade at the root
and run along the top and bottom surfaces of the fairing, as shown in Figure 4-1.
This is
in contrast to the flex being located in the web of the Rodgers blade, as shown in Figure 2-5.
This allows the electrode tabs to protrude straight out from the spar laminate, eliminating the
need for a sharp bend in the tabs. It also removes the connections from the very cramped area
of the web and allows for greater separation of connections.
The design of the actuator packs and flexible circuit, and the blade manufacturing procedure
78
that follows in Section 4.5, will be based on the design and manufacturing of blade section 4.0.
The name of this blade section was chosen to follow the naming convention established in
Rodgers' work [Rodgers, 1998]. Blade section 4.0 is a 21" long section of a blade that is built to
test the design and manufacturing changes made to the Rodgers blade design and manufacturing
procedure.
Chapter 5 will discuss the design and manufacturing of the Active Material Rotor
(AMR) blade, a full model-scale integrally actuated blade.
The AMR blade is built using the
design and manufacturing techniques developed in this chapter.
4.3
Actuator Pack Leadout Design
Blade actuator packs for twist actuation are 450 orientation packs, as seen in Figure 4-2. Each
actuator pack has two leadouts which are used to connect to the actuator pack.
The pack
width is determined by the chordwise extent of the active material in the spar of the blade.
The spanwise length is limited by AFC manufacturing limitations [Pizzocchero, 1999] and blade
robustness requirements.
If an AFC were manufactured to cover the entire active length of the
blade, a failure in the pack would cause the entire ply to be lost. Having multiple AFC packs
per ply provides redundancy in actuation.
The actuator leadouts are designed to protrude from the spar laminate during and after the
spar cure.
See Figure 4-10.
This prevents bending of the electrode tabs.
Bending the tabs,
as was done in the Rodgers blade, creates possible stress concentrations at the bend. Also, the
screen printed silver ink is believed to be more susceptible to cracking at sharp bends than the
copper electrodes.
The leadout tabs are shown in Figures 4-2 and 4-3. A 1" long, 1/4" wide copper strip is
bonded to the silver ink electrode with conductive epoxy. The tab is trimmed to leave 1/8"
of Kapton on each side of the copper strip to reinforce the bond. The Kapton is bonded with
AFC matrix material, which is stronger than the conductive epoxy, helping to prevent peeling
apart of the upper and lower electrode tabs. This holds the copper strip securely in place.
This processing performed by the AFC pack vendor1 .
the solder pads on the flexible circuit.
The copper strip is designed to reach
Figure 4-1 shows a schematic of the electrode tabs
'Continuum Control Corporation, Billerica, MA., 978-670-4910.
79
II
6.29 (AMR), 5.77 (Section 4.0)
-
0
CDS
Copper Strip
Electrode Tabs
(Electrode Leadouts)
Figure 4-2: Typical blade actuator pack. Dimensions are shown both for the AMR blade
design and the blade section 4.0 design. Two leadouts power the actuator. Each leadout has
a copper strip extending from it. This copper strip is soldered to the flexible circuit.
Electude Pattern
0.15"
Kapton trim
0.125"
Copper
Strip
-0.125"
- 0.250"
Figure 4-3: Electrode leadout tab detail.
80
protruding from the spar and running aft to the flexible circuit solder pad. The manufacturing
of the connections will be described in detail in Section 4.5.
It was determined that to mount a copper strip securely to the tab, the necessary electrode
tab length was 0.20" [Pizzocchero, 1999]. This leaves the electrode tab protruding 0.15" from
the pack trim edge, as seen in Figures 4-2 and 4-3.
The spacing of the leadouts is chosen to maximize the distance between any two adjacent
leadouts on the trailing edge of the spar.
This reduces the risk of short circuits between the
leadouts due to small shifts of the packs during the spar cure.
4.4
4.4.1
Flexible Circuit Design
Motivation for Using a Flexible Circuit
Each actuator must be individually electrically accessible from the exterior of the blade. The
resistance and capacitance of individual packs have to be measurable after the blade is cured.
This requires two individual, dedicated wires per actuator packs.
Given that an active blade
incorporates several dozen actuator packs into the blade, many individual leads are necessary.
Using conventional wire for such a task would require a large, heavy wire bundle.
would have to withstand a voltage of 3KV. This would require relatively large wire.
Resist wire from Gore, Inc. 2 can withstand 4KV, and is 0.060" in diameter.
The wire
Corona
A bundle of
50 such wires is approximately 0.5" in diameter, too large and heavy to be incorporated in a
blade.
The forces exerted on the blade components by the spinning of the blades is enormous,
equalling approximately 2500 times the force of gravity at the tip, so even a small amount of
mass can create tremendous forces on the structure.
the 1/4 chord of the blade.
Also, the blades must be balanced about
Using a heavy wire bundle aft of the 1/4 chord would in turn
require a significant amount of weight to be placed at the leading edge to balance the blade.
To solve the bulk and weight problems associated with a wire bundle embedded in the blade,
a flexible circuit is used to provide power to the actuators.
Flexible circuits are thin and light
and can run many lines in close proximity to each other. Flexible circuits consist of copper
lines in an acrylic matrix, sandwiched between two plies of Kapton. The acrylic matrix in the
226 gauge, 4KV, Corona Resist Wire FO1A080, W. L. Gore and Associates, Inc., 1-888-914-4673
81
0.1 "x. I" solder pads
0.010" copper lines
Spaced 0.020" apart
Figure 4-4: Blowup of flexible circut, showing 0.010" lines and solder pads. The lines are spaced
0.020" apart, edge to edge.
flexible circuits has a breakdown voltage of 3KV over 0.001". Therefore, barring defects in the
circuit, two lines spaced apart a few thousandths of an inch in a flexible circuit should be able
to withstand the driving voltages [Allflex, 2000].
4.4.2
Flexible Circuit Design
To enable access to the connections between the packs and the flexible circuit, the flexible
circuit is routed onto the upper and lower surfaces of the fairing. This presents a significant
design change from the first generation blade, in which the flexible circuit ran along the web of
the blade. The new design allows for direct heating of solder joints, as well as inspection and
possible repair to the connections.
The flexible circuit consists of several layers.
Blade section 4.0 has two flexible circuit
layers. The Rodgers blade used 6 layers, and the AMR blade to be discussed in Chapter 5 will
incorporate 4 layers. The number of flexible circuit layers is driven by the number of actuators
embedded in the blade.
Since line width and line spacing requirements in the flexible circuit
must be met, having more lines in the blade will increase the number of flexible circuit layers
needed to carry all the lines.
2
The lines are designed to be 0.010" wide, and made of loz/ft
copper. These lines have a current carrying capability of 0.5A, well above the current draw of
82
a single pack (20mA, see Section 3.7.5), The large lines reduce the risk of resistive heating in
circuit lines. It also reduces the effect of small defects in the lines and increases the strength of
the circuit [Allflex, 1998].
In the first generation blade, some defects in the flexible circuit caused short circuits.
To
reduce the risk of short circuits due to defects in the circuit, the separation of the lines was
increased form 0.015" (in the first generation blade circuit, [Rodgers, 19983) to 0.020" edge to
edge. Figure 4-4 shows the design of the flexible circuit used in blade section 4.0.
The flexible circuit enters the blade at the heel of the spar, near the root of the blade.
The circuit then runs along the web of the blade.
The upper and lower inboard circuit layers
transition to the fairing surfaces immediately inboard of the first actuator electrode tab. These
two layers then run along the fairing surfaces, where the connections to the actuator packs are
made.
To keep weight as far forward as possible in blades requiring more than two layers of
flexible circuit, any outboard flexible circuit layers run along the web until the transition from
the web to the fairing surface occurs outboard of the outboard-most connection on the inboard
layers.
See Figure 4-1.
The connections are made between copper strips exiting the electrode tabs and the respective
flexible circuit solder pads.
4.5
Blade Manufacturing Procedure
This section will describe the manufacturing procedure used to build the blades with the new
power bus design.
Parts of the manufacturing that are similar to the first generation blade
manufacturing procedure will be left out.
Therefore, such steps as foam shaping or ply layup
in the spar will be omitted. This information has been published [Rodgers, 1998].
4.5.1
Blade Section 4.0 Overview
Blade section 4.0 is built to develop the blade manufacturing procedure using the redesigned
components discussed in Sections 4.3 and 4.4. The crossectional geometry of the blade section
is that of the CH-47D model scale blade. The chord of the blade is 5.388" and the thickness is
0.647".
The blade section design and layup is shown in Figure 4-5. Six AFCs are embedded in
83
.........
..
...........
...............
SPI: E-glaso 0 dog, 0.0045*
SP2: S-glass 0 45 dog, 0.009"
LI
SP3: S-glass @ 0 dog, 0.009"
4-
SP4: S-gloss 0 -45 dog, 0.009"
SPS: S-glss @ 0 dog, 0.009"
001
SP6: S-glss @ 45 dog, 0.000"
SP7: S-gloss @ 0 dog, 0.009"
SPS: E-glass @ 0 dog, 0.0046"
B-
0LL
SPI: E-glass @ 0 dog, 0.0045"
SP2: AFC @45 dog, 0.0065"
SP3: S-glass @ 0 dog, 0.009"
SP4: AFC 0 -45 dog, 0.0085"
4-
SP5: S-glass @ 0 dog, 0.00o"
AFCs
SP6: AFC @ 45 dog, 0.0085"
20.63"
SP7 S-glass @ 0 dog, 0.009"
@0 dog, 0.0045"
SP8: E-glass
0-
Active Area
Web
(3 plies, 450E-glass)
Fairng Skin
(1 ply 0* E-glass)
1.92"
2.05"
5.388"
Figure 4-5: Drawing of blade section 4.0. The layup is shown for the active and non-active
areas. The non-active area incorporates S-glass filler plies to substitute for the AFCs. Tip
fittings are embedded in the ends of the spar to allow for fixturing. Detailed drawings are
provided in Appendix B.
84
Figure 4-6: Electrodes protrude from the spar laminate during layup. As the active plies are
laid into the spar, the packs will tend to lift off the inner plies due to high actuator stiffness.
the section. These AFCs are 5-mil fiber AFCs, with silver ink electrodes. The layup of blade
section 4.0 is torsionally soft. The layup eliminates all 450 passive plies from the laminate. The
passive reinforcement for the AFCs is therefore removed, leaving the packs more susceptible to
damage, but allowing more blade twist.
Detailed drawings of the blade section and blade components are presented in Appendix B.
The manufacturing procedure that will be used to build blade section 4 will be similar to the
procedure used to manufacture the full blade.
Since the blade section does not have a root,
the root manufacturing procedure will not be presented.
The root manufacturing procedure
can be found in [Rodgers, 1998].
4.5.2
Spar Layup
The foam core and the composite plies are prepared based on the manufacturing procedure
published by Rodgers [Rodgers, 1998].
After the web plies and the innermost passive plies
have been laid up, the AFCs are laid into the spar. Prior to layup, the AFCs are cleaned with
acetone and methanol to remove any residue from manufacturing or handling.
removes any residue left from the acetone.
The methanol
The electrode tabs are left protruding from the
spar. Hand pressure is applied to the AFCs to make them stick to the passive material around
85
Figure 4-7: GNPT tape is applied to the electrode tab and copper strip. This prevents resin
from running onto the strip, and prevents the tab from bonding to the mold during the cure.
The copper strip can be folded onto itself for ease of handling. The tape can be applied either
before layup of the packs in the spar, or immediately prior to the spar cure.
them. As the successive plies of AFCs are laid into the spar, the laminate will tend to lift up
from the innermost passive plies due to the stiffness of the AFCs. This can be seen in the lower
surface in Figure 4-6. Care must be taken to keep pressure on the laminate to ensure that the
plies do not move out of position.
4.5.3
Electrode Tab Protection
During the spar cure it is imperative that no epoxy leaks onto the copper strips exiting the
actuator leadouts. This would prevent soldering to the flexible circuit. Removal of epoxy from
the copper surface is difficult and may lead to damage to the leadout. To prevent any damage
or epoxy leakage onto the copper strip, the copper strip is covered with GNPT tape, as shown
in Figure 4-7. The GNPT also prevents the electrode tab from being bonded to the hardback
or the mold during the blade cure. The copper strip can be bent onto itself to make the spar
easier to handle and less prone to damage prior to the cure. The copper is malleable enough
that even a sharp bend in the strip will not cause damage.
4.5.4
Spar Cure
A hardback is used during the spar cure to apply pressure to the web of the blade. A hardback
is essentially a part of the mold.
It is used to fill the portion of the mold where the fairing
86
.....
..........
.......
.......
.................
Cutting Plane
Aluminum Tape Filler
Hardback Cure
411
31
Web
Spar Cure
Spar Laminate
Figure 4-8: Hardback during hardback cure and during spar cure. The number of S-glass plies
necessary to fill the mold at each chordwise location is shown in the hardback picture. The
aluminum fill is 0.050" thick. This makes a depression in the hardback that accommodates
the electrode tabs during the spar cure. During the spar cure, the hardback applies pressure
to the web of the blade.
Figure 4-9: The hardback is mounted in the mold, and the spar is slid into place. This picture
shows the electrode tabs not covered in GNPT. However, the tabs are covered in GNPT before
the spar is inserted into the mold. See Figure 4-7.
87
would normally be located. The hardback is necessary when making the D-spar since the web
must be compressed during the spar cure.
The hardback is manufactured by laying S-glass 3 into the mold.
The number of layers of
S-glass that are laid into each chordwise location in the mold is predetermined to ensure that
the correct amount of S-glass is used.
See Figure 4-8.
The mold must be completely filled,
but not overfilled, since overfilling would prevent proper closing of the mold.
To create the depressions in the hardback, aluminum tape is laid into the mold immediately
aft of the web line.
The aluminum tape is stacked to a height of 0.050",
approximately the
thickness of the spar laminate, to create a depression which will ensure that the entire web of
the blade is pressurized and will allow enough room to accommodate the electrode tabs and
copper strips.
See Figure 4-8.
After the hardback is cured in the mold, it is cut using a water jet cutter. A sanding wheel
is used to do the detail work on the web portion of the hardback, specifically the blade pretwist
that must be followed by the web of the blade.
The hardback is held in place in the mold using two screws.
The holes for the screws are
drilled through the hardback, into the mold. This is done before the hardback is removed from
the mold after the cure.
This ensures that the hardback is secured in the correct location of
the mold.
Once the hardback is complete, all hardback surfaces that will be in contact with the blade
are covered in GNPT tape. This prevents the blade spar and electrode tabs from bonding to
the hardback during the cure.
The spar is laid into the mold, with the electrode tabs protruding into the depression in
the hardback.
The mold is then closed and the spar is cured.
Details on the spar cure and
mold closing can be found in [Rodgers, 1998].
Following the cure, the mold is opened carefully,
and the spar is removed from the hardback.
Care is taken to pry the electrode tabs off the
hardback before any force is applied to the spar to cause it to separate from the hardback.
If
the electrode tabs are bonded to the hardback or the mold, applying force to the spar to pry it
off the hardback, or out of the mold, can cause the electrode tabs to be damaged or torn off.
The GNPT protective tape is removed from the electrode tabs using tweezers.
3
SP381-UNI-S29 284BW 33RC, 3M Aerospace Adhesives, 1-800-235-2376.
88
The tabs
..............
..........
Figure 4-10: The electrode tabs protrude from the spar. After the GNPT protective covering
is removed, the electrode tabs are pre-tinned to allow for good heat conduction into the solder
joint to the flexible circuit.
are straightened, and any resin that leaked through the GNPT protection is scraped off the
copper strips. To scrape epoxy off the copper strips, the strip is laid on a hard, smooth surface,
and the sharp point of a razor blade is used to remove the epoxy. The cured spar is shown in
Figure 4-10.
4.5.5
Fairing Assembly
The blade manufacturing continues with the assembly of the fairing. The fairing foam is shaped
and assembled [Rodgers, 1998].
fairing core assembly.
The flexible circuit layers are cleaned and mounted to the
The upper layer is laid onto the foam first, followed by the lower
layer. To keep the circuit layers in place, they are tacked down with 5-minute epoxy wherever
4
necessary, usually in 3 or 4 spots. Alternatively, flash tape can be used to temporarily secure
the flexible circuit to the fairing foam.
The flexible circuits run along the web as far outboard as possible to keep weight forward in
4Flash tape, Richmond Aircraft Products, Norwalk, CA.
89
.....................
.. . ......................
Electrode Tabs
Flexible Circuit
Depressions in
Fairing Core
Fairing Core
Figure 4-11: AMR blade fairing assembly.
the blade. The circuits transition to the fairing surfaces immediately inboard of the connections
at
that they will connect to. See Figure 4-1. A small radius is made in the flexible circuit layers
blade
the transition from the web to the fairing. This reduces the risk of breakage during the
cure. See Figure 4-12.
A foam insert is placed on top of the radiused circuit layer to provide
back pressure to the fairing skin during the fairing cure.
is inserted between the flexible circuit layers on the web.
See Figure 4-14.
Film adhesive5
This ensures a good bond between
flexible
the flexible circuit layers in the web. No film adhesive should be inserted between the
the
circuits and the foam on the fairing surface, since a heat barrier must be inserted between
flexible circuit and the fairing foam when the electrical connections are made.
Film adhesive
would prevent this. More detail will be given later in this section.
For the fairing assembly to fit snugly against the web of the blade, it is necessary to make
from
small depressions in the fairing foam to accommodate the electrode tabs that protrude
from the
the spar. These depressions are either made by removing small amounts of material
is repeated
foam, or by using a blunt object to make a depression in the foam. This procedure
for each electrode tab location.
See Figure 4-11.
Once a good fit is achieved between the spar and the fairing assembly, the fairing assembly
is held in place with flash tape and the connections are made.
5AF-163-2U03 film adhesive, 3M, Austin, TX, 800-235-2376
90
The connection between the
Figure 4-12: Flexible circuit laid onto the fairing foam. Radius in the transition area decreases
chance of breakage during the fairing cure. The marks on the foam indicate location of foam
depressions.
Figure 4-13: Blade section 4.0 with connections to flexible circuit made. Copper strip excess is
trimmed off. Depressions that accommodate the electrode tabs can be seen. These depressions
are filled with either foam or foaming adhesive prior to the fairing cure.
91
Foam Insert
Flexible Circuit
Figure 4-14: Crossection of flexible circuit layer transition from the web to the fairing surface.
Corresponds to location A in Figure 4-5. Not to scale.
Foam or~
Foaming Achesive
Insert
Fairig Skn
Electrode.!
A
Solder
oint
Flexible
Circuit
-
Depression in Fairing
Foam
-e d
Figure 4-15: A foam insert or foaming adhesive is placed on top of the electrode tab to provide
back pressure to the fairing skin. Corresponds to location B in Figure 4-5. Not to scale.
92
copper strip mounted on the electrode tabs and the solder pads on the flexible circuit are made
with a solder joint, using a regular lead/tin solder with a melting point of 361'F. The copper
strip is pre-tinned on both sides.
The solder pads on the flexible circuit are supplied pre-
tinned, but extra solder can be applied to the solder pads if the supplied tinning is sparse. To
make the connection, the copper strip is laid on top of the solder pad, and heat is applied to
the top surface of the copper strip.
The joint is heated for a few seconds until solder is seen
squeezing out from under the copper strip. A heat barrier consisting of a few plies of pre-cured
E-glass is inserted between the flexible circuit layer and the fairing foam to prevent melting or
deformation of the fairing foam underneath the solder pad. All connections should be visually
inspected.
Electrical properties of the actuator packs should be measured through the flexible
circuit to ensure that the connections to the packs have been made properly. Figure 4-13 shows
the completed blade section 4.0 connections.
After all the connections have been made, film adhesive is inserted between the flexible
circuit layers and the fairing foam. This is to ensure a good bond between the circuits and the
foam.
The E-glass fairing skin bonds to the upper surface of the flexible circuit, and no film
adhesive is necessary between the flexible circuit and the fairing skin.
The area above the electrode tab needs to be filled to provide backpressure to the fairing
skin.
This is done either by inserting a piece of foam of the appropriate thickness or by using
some foaming adhesive. The foam insert can be made by crushing a small piece of foam down
to the correct thickness. In blade section 4.0, 0.010" thick inserts are used for the electrode tabs
of the AFCs in spar ply 2 (SP2), the active ply closest to the outer mold line.
See Figure 4-5.
0.025" thick inserts are used on the SP 4 tabs, and 0.040" inserts are used on the SP6 tabs.
The inserts will expand slightly to give backpressure to the fairing skin.
See Figure 4-15.
The foaming adhesive may be a better solution as it expands to fill all empty spaces around
the electrode tab and the copper strip.
However, caution should be used when applying the
foaming adhesive to estimate the filling volume of the adhesive correctly, so that the expansion
of the adhesive does not "overfill " the cavity, damaging the electrode tab or deforming the
foam core.
The fairing skin and any trailing edge details are then laid up.
assembled, it is inserted in the mold and cured.
93
Once the fairing is fully
A trough should be machined in the mold
to accommodate the flexible circuit where it exits the root of the blade [Rodgers, 1998].
Care
should be taken to avoid pinching of the circuit when closing the mold.
Connectors are installed on the inboard ends of the flexible circuit.
potted to ensure they can withstand high voltage.
These connectors are
The connector installation and potting
procedure is described in Appendix D.
4.6
Blade Section 4.0 Manufacturing Results
The completed blade section is shown in Figure 4-16. The flexible circuit and all 12 connections
that were made survived the blade cure. The backpressure on the fairing skin was adequate,
with only minimal wrinkling of the skin around the electrode tabs.
Using foaming adhesive
rather than foam inserts to fill the cavity around the electrode tabs will prevent any wrinkling.
The mold proved to be difficult to close completely during the spar cure.
The spar foam
had been shaped for a thinner spar laminate than was used in blade section 4.0.
The extra
core material made it difficult to close the mold completely, and very high clamping forces were
used.
This may have applied very large through thickness pressure to the spar laminate and
the AFCs. Even with high clamping pressure, the mold was not able to be closed completely.
The mold halves were separated by approximately 0.030" during the spar cure.
4.7
Actuation Testing Setup and Procedure
The test setup consists of 3 main components:
and data acquisition system.
laser displacement sensors, power amplifier,
Detailed descriptions of the test setup have been published by
Rodgers, and will only be discussed briefly here [Rodger, 1998].
The amplifiers that are used to drive the blade section, and the blade that will be described
in Chapter 5, are based on a pair of Yorkville AudioPro 34006 audio amplifiers. These amplifiers
are low voltage amplifiers, with a peak voltage of +/- 144V.
Therefore, to boost the voltage
capability of the amplifiers, each channel is run through a 25:1 transformer. Four channels are
available in the setup. Voltage and current protection are incorporated on the high voltage side
6
Yorkville AudioPro3400, Pickering, Ontario, Canada.
94
Figure 4-16: Completed blade section 4.0. Actuator packs can be seen in the spar, and the
flexible circuit can be seen clearly in the fairing. The flexible circuit and instrumentation wires
exit the blade at the inboard end.
95
of the amplifier. Varistors 7 are used to limit the maximum voltage output.
Electronic circuit
breakers, with a maximum current capability of 0.250A per channel, are used to protect the
test article from high currents that can occur during short circuit situations [Rodgers, 1998].
The data acquisition system consists of a National Instruments data acquisition system.
The system has 24 analog input channels, and 24 strain gage conditioning modules.
The
collected data is stored on a Pentium II based computer, and post-processed using MATLAB 8.
The displacement sensors used to gather actuation data are a pair of Keyence 9 LB-12 laser
displacement sensors.
These lasers have a range of 20mm, centered at a distance of 40mm.
The resolution of the sensors is 50pm at a sampling rate of 3kHz.
The laser sensors have a
voltage output, with a calibration factor that is set to approximately 600V/m.
The blade is actuated in twist by driving the +45' and -45' packs 1800 out of phase. As
a result, when one set of packs is extending, the other is contracting.
deformation of the blade structure.
described above.
This creates a twisting
The twist is measured using the displacement sensors
Figure 4-18 shows the method used for gathering twist data.
The twist
angle, 0, is found by,
0 = atan(
L
),
(4.1)
where dl is the displacement measured on laser 1, as seen in Figure 4-18, d 2 is the displacement
on laser 2, and L is the distance between the lasers.
During actuation testing, the root of the blade is fixed and the tip is free.
The test rig
that is used can be seen in Figure 5-16.
When driving the blade, care is taken to avoid sharp steps in the drive signal.
The drive
signal is dialed up slowly to avoid driving the blade with the high frequency content of the step
that would result from a sudden power-on.
The high frequency content or a step or a sudden
spike could possibly damage the actuator packs.
7
Harris Metal Oxide Varistors, 570-V575LA80B, Mouser Electronics, Randolph, NJ. 800-346-6873.
"MATLAB, The MathWorks, Waltham, MA.
9
Keyence LB-12 laser displacement sensors. Woodcliff Lake, NJ. 201-930-0100.
96
Figure 4-17: Active area of completed blade section 4.0. The connections to the flexible circuit
are visible, along with the filler material placed over the electrode tabs. The backpressure on
the fairing skin was adequate, with minimal wrinkling around the elelctrode tabs.
d2
di
L,
CI)
CO
C')
-j
-
Figure 4-18: Laser displacement sensors are used to measure the actuation level of the blade.
Eq. 4.1 gives the formula for calculationg 0, the blade twist angle.
97
-.....................
..........................
........
-
0.3
0.25
0.2
~0. 15
0.1
0.052p.D
0
-1.
-
-
~520
-
O.$
Voltage (kV)
Figure 4-19: Blade section 4.0 actuation.
distribution along the span.
4.8
Laser vibrometer data shows graphically the twist
Blade Section 4.0 Testing
Blade section 4.0 achieved a 1.64/m twist rate at 215OVpp, 0VDC. The model prediction, using
a Boeing FEM model, was 2.0*/m [Schmidt, 2000].
Figure 4-19 shows actuation data for blade
section 4. A laser vibrometer' 0 was used to provide visualization data of the twist distribution
of the blade section.
The blade section twist did not meet model predictions. Since all available actuator packs
were used in the blade section, individual pack testing such as -that described in Chapter 3
was not possible. The post-lamination behavior of the actuator packs could therefore not be
tested using controlled laminations.
However, based on the data presented in Section 3.5, it
is possible that the packs had a high actuation reduction as a result of lamination. The blade
section 4.0 packs were 5-mil fiber AFCs, and had an average actuation level of 1083pE. If these
packs behaved like the packs that were tested in Chapter 3, it is likely that the blade section
4.0 packs would have had a large reduction in actuation due to lamination.
However, since no
spare packs were manufactured, this hypothesis can not be verified.
Another possible explanation for the lower than expected performance is pack damage. As
explained in Section 4.6, the core of the spar was designed for a thinner laminate than the one
10
Scanning Laser Vibromter, Model PSV-300-F, Polytec PI, Inc. Tustin, CA
98
that was used.
It therefore proved difficult to close the mold without using excessive force.
This could have created tremendous through thickness pressure in the mold, possibly damaging
the actuators.
Also, the mold was not preheated prior to closing.
resin at room temperature is quite high,
Since the viscosity of ply
significant shear forces may have been introduced
into the laminate, possibly damaging the actuators [Weems, 1998].
The lack of 450 plies in the
laminate may have exacerbated the damage to the AFCs.
4.9
Summary
This chapter discussed the development of a new power bus design and manufacturing procedure
to improve blade manufacturability and robustness. The blade sections fulfilled the goals that
were set. The blade section manufacturing showed that the blades are buildable using the new
manufacturing procedure and the new actuator packs.
The new design took into account the
problems faced in earlier blades and the recent changes in actuator material. The manufacturing
was made simpler by removing the connections from the cramped and hidden web, and bringing
them to the surfaces of the fairing. The improved manufacturing procedure reduces the chance
of damage to the packs and of short circuits in the flex circuit and connection areas.
99
100
Chapter 5
AMR Blade Design and Testing
The MIT/Boeing Active Materials Rotor (AMR) Blade is the first full model scale blade being
built with the manufacturing procedure developed in this thesis, and with the new 0.010" fiber
AFC actuator packs. Four fully active blades are being manufactured by Advanced Technologies, Incl (ATI).
The blades will be windtunnel tested at Boeing Helicopters, Philadelphia.
Two additional blades are being built as well.
One is a fatigue blade, which will be fatigue
tested following the windtunnel testing. The other is a process validation blade, known as the
process blade. This blade was built by ATI to validate the manufacturing process and to give
manufacturing practice prior to building the flight blades.
The process blade was half active,
incorporating only the inboard 12 actuator packs. The process blade was bench tested to verify
the integrity of the actuators and connections.
This chapter will describe the blade design and present the test result of the process blade.
Most structural design work was performed by Boeing Helicopters, and will not be discussed in
detail. The design work performed by MIT on the power bus and the AFCs will be presented.
The design of the flexible circuit external to the blade and the power distribution board will be
discussed. Finally, bench test results of the process blade will be presented.
actuation data, current draw data, and heating data.
'Advanced Technologies, Inc., Newport News, VA., 757-873-3017.
101
This will include
5.1
5.1.1
AMR Design
Blade Geometry and Layup
The AMR blade is a 1/6 scale, Mach scaled advanced CH-47 blade. The blade, like the CH-47D
blade, consists of a D-spar and an aerodynamic fairing. It is considerably wider than the first
generation blade (6.10" vs. 5.39"), and slightly thinner (0.61" vs. 0.65").
It has a swept tip
that tapers from 6.10" chord to 3.66" chord from 0.92R to 1.OOR. A drawing of the blade is
shown in Figure 5-1 [Bussom, 2000].
The AFCs run from 0.27R to 0.92R, giving a total active length of 39.0".
Two plies of
AFCs are in the top and bottom spar laminates. Each ply consists of 6 AFC actuators.
The
AFC geometry is presented in Section 5.1.2. The laminate consists of both 00 and 45' passive
plies.
The 0' plies are S-glass plies, and provide axial strength and bending stiffness.
The
45' plies are E-glass cloth, and provide reinforcement to the AFCs in the fiber direction. The
active span layup is shown in Figure 5-2.
5.1.2
Actuator Packs
The AMR blade was the first blade to be built with actuator packs manufactured with 0.010"
diameter fibers. The change from 0.005" diameter fibers to 0.010" diameter fibers allowed for
fewer active plies, significantly reducing manufacturing difficulty and blade complexity. Fewer
active plies meant fewer packs, which in turn meant fewer electrical connections that had to be
made.
Figure 5-3 shows the design of the AMR blade actuator pack and electrode. The pack is
6.29" long and 1.77" wide. The pack length allows 6 packs to the be incorporated in a ply, with
0.29" gap between packs.
This gap is incorporated between the packs to prevent any short
circuits from occurring between packs. This gap is filled with S-glass in the spar layup. The
pack width was limited because of high curvatures at the nose of the blade. The high curvature
could possibly damage the fibers in the pack, and the nose area is therefore avoided.
The pack has two electrode tabs on the trailing edge of the pack. The tabs are an extension
of the AFC rails, and are used to supply power to the actuator.
The tab near the end of the
pack is designated as the positive tab, and the tab near the middle of the pack is designated as
102
55.20
16.20
0.92R
0.27R
9.026R
A.1P
39.
6.29"--
I
"
60.00
1.0OR
0.29"
_%
I
041
X, .
X
I VI "I,
[.
6.10"
I'
3.66"
I
I- I I
50.98"
~C~9
Active Area
0.61.
1.77"
6.10"
Figure 5-1: Active Material Rotor (AMR) blade drawing.
0.0045"
E120 @45 deg
0.0135"
AFC @45 deg
0.009"
S2 @ 0 deg
S2 @0 deg
0.009"
AFC @-45 deg
0.01 35"
0.0045"
E120 @45 deg
0.009"
S2 @0 deg
0.0035"
S2 (thin) 0 0 deg
Figure 5-2: Typical AMR blade layup, representative of the active section from 0.34R to 0.65R.
The AMR layup has two plies of AFCs. The 450 E120 provide reinforcement to the AFCs in
the fiber direction. The 0' S2 provides flap stiffness and axial strength.
103
1. 77"
0. 158"
3.000"
0.15"
6.29"-
0.125"-
0.125"
0.250"
m
CD
Figure 5-3: Electrode pattern and trim line of the AMR AFC actuator packs.
the negative tab. The distance between the tabs was chosen to maximize the distance between
the electrode tabs in the blade.
The packs may shift during the cure, and having maximum
spacing between leadouts minimizes the chance of overlap between tabs.
The outline in Figure 5-3 shows the line along which the pack is trimmed. The pack is
trimmed 0.05" from the trailing edge of the actuator pack electrode and 0.03" from the other
3 edges of the electrode. The trim distance on the trailing edge is larger than on the other
edges of the packs to prevent arcing between packs on the trailing edge of the spar laminate.
If the packs were trimmed to the electrode pattern, and the edge of the pack was aligned with
the web of the blade, arcing could occur between packs in the +45'
and -45* plies.
Leaving
extra trim material on the trailing edge pushes the electrode pattern forward from the web,
preventing arcing.
The electrode rails are 0.045" wide, and the spacing of the electrode fingers is 0.045". Each
electrode finger is 0.0075" wide. This geometry is typical of the standard electrode pitch used
104
0.44"1
0.51"
0.1xO.1" solder pad
0.77"1
\
Transition region
Figure 5-4: Internal portion of the flexible circuit, showing the transition region, the location
in the circuit where the transition from the web to the fairing occurs. A more detailed drawing
of the flexible circuit, including location of the solder pads, is provided in Appendix C.
by the AFC vendor 2 .
5.1.3
Internal Flexible Circuit Details
The flexible circuit for the AMR blade was designed as described in Chapter 4. Figure 5-4 shows
the dimensions of the flexible circuit that runs internal to the blade. The section of circuit that
runs on the web is 0.44" wide and is designed to run between the protruding electrode tabs on
the web.
See Figure 5-5.
The transition area from the web to the fairing surface consists of
an offset in the circuit of 0.77". This "step" allows the circuit to exit the web of the blade,
making a 0.08" radius to the surface of the fairing.
The section of circuit that runs on the
fairing surface is wider than the portion that runs on the web, 0.51". This larger width is due
to the 0.1" square solder pads that have to be accommodated.
The circuit on the fairing surface is designed to run so that the leading edge is 0.25" behind
the web. This accommodates the electrode tabs and allows the copper strip to transition from
the tab to the solder pad.
It also keeps the circuit as far forward as possible to reduce the
amount of leading edge weights required to balance the blade about the quarter chord. Each
flexible circuit layer weighs approximately 1g. per foot.
Although this is fairly light, it still
must be taken into account when balancing the blade, and even small amounts of added weight
can cause large increases in hub loads due to the tremendous centrifugal force loads on the
blade.
2Continuum
Control Corporation, Billerica, MA., 757-873-3017.
105
................
....
.......................
.. .....
Spar Laminate
Section B
0.08" radius
Flexible Circuit
0.44"
Fairing
Section A
Tab Exit Location
Figure 5-5: AMR blade crossection showing the flexible circuit, The two figures correspond to
crossections A and B as seen in Figure 5-1. The portion of the circuit that runs on the web
fits between the electrode tabs, and allows a 0.08" radius to be made to the fairing surfaces.
During testing of the flexible circuit, it was noted that short circuits were occurring between
the inboard-most solder pad, and the line for the second solder pad. Upon inspection under a
microscope, it was noticed that the cutout for the solder pad overlapped the second line on the
flexible circuit.
See Figure 5-6.
This caused an arching path between the solder pad and the
line. In future flexible circuit designs, the solder pad should be removed farther from the line
so that the cutout does not overlap any other lines.
5.1.4
External Circuit Design
The inboard portion of the flexible circuit consists of the portion of the circuit that runs from
the exit point from the blade to the hub of the rotor. The design of this portion needed to
take into account routing around the lead-lag and flapping pins, the routing and attachment on
the pitch housing and instrumentation can, and the interface to the power distribution board.
Also, airloads in hover and in forward flight had to be taken into account, and the possibility of
replacement of the inboard portion of the flexible circuit in case of damage had to be considered.
Damage to the flexible circuit is most likely to occur in the area that runs over the flapping
106
.... ....
...........
Figure 5-6: The cutout for th e r ad in the flexible circuit was too large, overlapping the
adjacent line. This 1e&fd-sh@(C$0C1ts between the solder pad and the line.
hinge. The combination of pitching and flapping motions in this region can cause significant
shearing and kinking in the circuit, which may lead to breakage. As a result, the ability to
replace this region of flex circuit is desirable. A "flareout" is incorporated in the flexible circuit
to allow access to the lines in case of breakage across the flapping hinge.
This flareout allows
soldering of wires or mounting of a connector to the flexible circuit layers.
The flareout is
placed inboard of the lead-lag hinge since there is less risk that the flexible circuit would tear
across this hinge than across the flapping hinge.
To avoid excessive forces on the circuit as a result of pitch link Hitions, a large loop must be
formed over the flapping hinge to allow the circuit to move freely without kinking or shearing.
The loop size was determined through mocking up the hub and determining the size loop that
would prevent snapthrough of the loop and minimize shearing forces on the circuit.
The loop around the flapping hinge must be protected from the airstream.
As a result a
skirt is placed around the instrumentation can and the inboard portion of the pitch housing.
This skirt is meant to shield the flexible circuit loop from airloads during hover testing and
forward flight testing.
To facilitate attachment of the flexible circuit to the pitch housing and the instrumentation
can, several "ears" are incorporated into the circuit to allow for tie down points.
can be seen in Figure 5-8.
107
These ears
.. ...
..
.....
........
.................
...........
Power Distribution Board
Instrumentation Can
T e Down Points
Flareout
/-- I
Flap Hinge
Blade
Lead-Lag Hinge
Pitch Housing
Figure 5-7: Rotor hub showing the flexible circuit routing.
0.50
0.25"
*
0.50"
0.25"
BS5.0
G10 Reinforcement Pad
BS7.5
BS10.0
Drill Through
for 8-32 screw
S3.5" 3.8"
Figure 5-8: External portion of the AMR flexible circuit. The flareout between BS5.0 and
BS7.5 allows for connection to the lines in case replacement of the inboard portion of the
circuit is needed. The offset at BS5.0 is designed to transition the flexible circuit from the
back of the pitch housing to the top.
108
. ...
4 Leads
.
.................
In
Power
Distribution
Board
T
ld
I
48 Leads Out
Rotating Frame
Frame
-Fixed
Top Surface
Bottom Surface
Channel I
Amplifr
Channel 2
Amplffier
Figure 5-9: The power distribution board distributes power to the 48 lines running into the
blade. The schematic shown is for each blade in the rotor.
Figure 5-10: Power distribution board with flexible circuits.
109
5.1.5
Power Distribution Board
Two power channels are available per blade.
hot wire.
Each power channel consists of a ground and a
The first channel corresponds to the +45' plies and the second corresponds to the
-450 plies. The flexible circuit has a total of 48 lines, and each of these lines has to be tied to
the correct channel wire. A circuit board was designed to distribute the power from the four
power leads running through the slipring to the individual lines on the flexible circuit.
The flexible circuit leadout consists of a potted Berg Clincher 3 connector with 12 leads,
spaced 0.1" apart. The leadouts are described in Appendix C. The power distribution circuit
board has four 12 pin female headers to accept the male connectors on the flexible circuit.
Female headers were chosen because they offer better safety.
When the flexible circuits are
not connected to the distribution board there is no exposed conductor that could pose a safety
threat.
To prevent short circuits or arching from occurring between the lines on the circuit board,
the board is potted with epoxy.
Shell 828 resin with 3223 curing agent 4 is used for potting.
This epoxy has low viscosity when mixed, and is able to fill any small voids on the board and
in the headers.
An air release agent is used to ensure that no air bubbles become trapped in
the epoxy 5
5.1.6
Options in Case of Flexible Circuit Failure
In the event that the flexible circuit were to break between the flareout on the pitch housing and
the instrumentation can, a wire bundle can be substituted for the flexible circuit. Using Gore
Corona Resist 6 wire, a wire bundle can be made that can run from the pitch housing flareout
to the instrumentation can.
The Gore wire is rated to 4KV and is 0.060" in diameter.
such wires creates a wire bundle that is approximately 0.5" in diameter.
48
One wire is soldered
to each individual flexible circuit line by removing an area of Kapton in the flareout with a
razor blade and soldering the wire directly to the copper.
allow them to function at high driving voltages.
3
The joints then must be potted to
Potting can be done with 5-minute epoxy
Berg Clincher connector. Newark Electronics P/N: 89F4613. Type: 65801-012.
Shell 828 resin, 3223 curing agent, Miller-Stephenson, Danbury, CT., 203-743-4447.
5Air release agent A530, BYK-Chemie, Wallingford, CT.
626 gauge, 4KV, Corona Resist Wire F01A080, W. L. Gore and Associates,
Inc., 888-914-4673.
4
110
.
since the area will not be exposed to high temperatures.
.
. ..........
. . . ........
....................
. . . . ......
The solder joints should be staggered
to maximize the distance between solder joints to further reduce the risk of short circuits. To
reduce the diameter of the wire bundle, some of the ground wires may be tied together on the
pitch housing, so that only a handful of ground wires run over the flapping hinge. All the hot
wires must still be run individually to allow electrical access to individual packs.
Substituting the broken flexible circuit portion with a similar portion of flexible circuit is
feasible, but not practical. Flexible circuit connectors would have to be mounted on the pitch
housing, an area that experiences significant centrifugal loads, making the attachment of the
connectors difficult.
Connectors between inboard and outboard flexible circuits that can not
be disconnected would simplify manufacturing and reliability of these connectors. Mating the
female and the male connectors prior to potting, and potting the connectors together eliminates
the need to keep connector receptacles clear of epoxy, a difficult task.
Manufacturing
5.2
The AMR blade was manufactured by Advanced Technologies Inc.(ATI) using the manufacturing techniques developed in this thesis. The first blade that was built was a tool-proofer blade,
or process blade. This blade served as a validation of the manufacturing process, and allowed
the manufacturer to gain experience with the manufacturing before building the wind tunnel
blades.
The process blade was half active, with only the inboard 12 packs operational.
Eight
electrically failed, non-operational actuator packs were incorporated in the outboard section of
the blade to better simulate the manufacturing procedure of a fully active blade.
"Dead" Packs
Live Packs
Figure 5-11: The process blade is only half active, with 8 "dead" packs outboard.
111
During manufacturing, one electrode tab broke off the spar heel.
was being removed from the mold following the spar cure.
This occurred as the spar
The tab was repaired by peeling
back Kapton on the tab to expose some copper, and soldering a new copper strip to the pack.
Tab breaking can be avoided by using a thin non-sticking film on the hardback surfaces that
are in contact with the tabs.
A total of 40 internal connections were made.
easily.
The connections were made quickly and
Continuity was achieved on all 40 connections, and was kept through the rest of the
manufacturing process.
The survival of all connections indicates that the solder connections
were robust, and that the electrode tabs and flexible circuits were not damaged during the
cure.
Foaming adhesive was inserted around the electrode tabs to fill in the void left by the
depression in the foam, as shown in Figure 4-15. The foaming adhesive flowed around the tabs
well, and did not appear to have deformed the electrode tabs, the copper strips, or the flexible
circuits.
A very smooth surface was achieved in the fairing skin around all the electrode tabs
and connections.
5.3
Bench Testing of AMR Process Blade
Testing is performed on the AMR process blade to validate the electrical connections under high
voltage and to demonstrate that the 10-mil fiber packs did not sustain any damage during manufacturing. Also, the process blade is the first AMR blade from which actuation measurements
could be collected.
The following is a list of testing performed on the blade.
" Twist actuation data is gathered from 500Vpp to 3000Vpp, at 500Vpp intervals.
1Hz,
22.8Hz, 45.6Hz, and 68.4Hz actuation frequency data is gathered at each voltage level.
This data is presented in Section 5.3.2.
" Transfer function data and scanning laser vibrometer 7 data is gathered.
This data is
presented in Appendix C. Transfer function data is gathered using the sine sweep method
developed by Rodgers [Rodgers, 1998].
7
Scanning Laser Vibromter, Model PSV-300-F, Polytec PI, Inc. Tustin, CA
112
Figure 5-12: Completed AMR blade.
the active area.
The flexible circuits can be seen running the length of
113
Figure 5-13: AMR blade root. The exit point of the flexible circuit can be seen. The flexible
circuits are wrapped in flash tape near the root for protection during testing. The instrumentation wire bundle is also shown.
114
Figure 5-14: AMR blade midspan detail. The transition of the flexible circuit from the web to
the fairing surface can be clearly seen.
Figure 5-15: AMR blade tip detail. The outboard-most AFC packs were not included in the
process blade. The empty connections can be seen on the flexible circuit.
115
"
Mechanical stiffness testing is performed to determine the bending stiffness and torsional
stiffness of the process blade. This data is presented in Section 5.3.3.
" Current draw testing is performed on the process blade. The current draw is determined
at 45.6Hz, at voltages ranging from 500Vpp to 3000Vpp, OVDC. The effect of adding a
DC offset to the drive signal is studied. 200VDC and 500VDC are applied to a 2600Vpp,
45.6Hz drive signal to determine the change in the current draw.
Also, DC offsets
ranging from 200VDC to 1000VDC are applied to a 4000Vpp, 22.8Hz drive signal. This
data is presented in Section 5.3.4.
" Heating testing is performed.
The blade is run at 2000Vpp, OVDC, at 22.8Hz, 45.6Hz,
68.4Hz, and 114Hz. The differences in heating at these frequencies are investigated. The
blade is also run at 2000Vpp, 2600Vpp, and 3000Vpp, OVDC at 22.8Hz to determine the
effect of varying voltage on the heating level in the blade. 200VDC and 500VDC offsets
are applied to a 2600Vpp, 22.8Hz drive signal to investigate the effect of a DC offset on
the heating level in the blade. This data is presented in Section 5.3.4.
To avoid early pack failures, the highest allowed blade voltage is incrementally increased.
Initial testing is performed at no higher than 100OVpp. Once actuation data has been gathered
at several frequencies, the maximum allowed voltage is increased to 2000Vpp. After actuation
testing at this voltage, transfer function and laser vibrometer data is gathered.
Following this
testing, the blade is driven to the full design voltage of 3000Vpp.
5.3.1
Testing Setup
The testing setup used to gather actuation data from the AMR process blade is similar to that
used for blade section 4.0, as described in Section 4.7.
seen in Figure 5-17.
The blade is clamped at the root, as
The blade exits the root clamp at BS10, 1" outboard of the root pin.
The tip of the blade is free.
The lasers displacement sensors and the testing rig are shown in
Figure 5-16.
Current measurements are taken from the current monitor on the high voltage amplifier.
The amplifier is discussed in Section 4.7.
The current monitor consists of a 1Q resistor in
116
The calibration factor of the current monitor is
series with the return path on the amplifier.
therefore 1V to 1A.
The resistance of the thermistor is measured using a digital ohm-meter
5.3.2
Actuation Testing Results
Blade models indicate that the process blade should yield 2.50 /m twist actuation at 3000Vpp,
OVDC [Weems, 1998].
This corresponds to a tip displacement of 1.1*, based on an 18" long
active area. Each pack has approximately 6" of active length, giving an active length of 18" for
the inboard half of the blade. The given actuation level is based on 1Hz actuation levels of the
embedded AFCs.
A 10% reduced AFC pack free actuation level is assumed in the modeling.
For a description of reduction factors, see Sections 3.4 and 3.6.
The actual average of the packs embedded in the process blade is 1339Pe at 3000Vpp, OVDC,
and 744wE at 2000Vpp, OVDC.
Based on these value, the predicted actuation level is 2.8' /m
at 3000Vpp, and 1.60 /m at 2000Vpp.
However, Chapter 3 showed that the free actuation can
not be used to predict laminated performance.
The high-field non-linearities of the material
are changed when the material is constrained.
Therefore, using the actual average of 1339[w
free actuation for the process blade packs may not yield valid predictions.
Using a reduced
free actuation for twist predictions at 3000Vpp will produce a more realistic prediction.
From
the data shown in Section 3.6, it can be assumed that a reduction factor should lie be between
10% and 25%.
Figure 5-18 shows the twist rate of the process blade at voltages ranging from 0 to 3000Vpp,
OVDC.
This data was collected with 11 of 12 packs operational.
The twist with 12 active
actuator packs can be extrapolated from this data, and is found to be 2.3* /m at the full design
voltage of 3000Vpp, 0VDC. This is within 10% of the predicted value based on 10% reduced
free strain values of the actuator packs.
25% free strain reduction.
Figure 5-18 also shows the predictions based on
25% is the reduction factor that was found in Section 5-18 to be
necessary to correct for the non-linear behavior of AFC packs with high free strain.
It is seen
that the actuation level of the process blade falls within the range of predictions incorporating
reduction factors of 10% to 25%, as given in Section 3.6.
8
Model DMM254,
Tektronix, Inc., Beaverton, OR.
117
Figure 5-19 shows actuation data at
Figure 5-16: At left, the blade can be seen mounted in the testing rig. This rig is used both for
actuation testing and for stiffness testing. Two laser displacement sensors are used to gather
deflection data.
Figure 5-17: Closeup of the root clamp of the testing rig. The root is clamped with a 3/8" bolt
that runs through the root pin hole in the blade. The blade exits the root clamp at BS10".
The clamp is bolted to a large aluminum base, as seen in Figure 5-16. The flexible circuit is
seen exiting the blade. Flash tape is applied to the flexible circuit for protection.
118
Only 8 of 12 packs are operational.
1Hz, 22.8Hz, 45.6Hz, and 68.4Hz.
As the frequency is
increased, the actuation level increases. This occurs because the driving frequency approaches
the first torsional natural frequency of 91Hz. For dynamic blade data, see Appendix C.
Twist versus span data is shown in Figure 5-20. The twist appears to vary linearly through
the active region. Figure 5-20 also shows a scanning laser vibrometer visualization of the twist
of the blade at 22.8Hz.
The twist can be seen varying linearly through the active region, the
inboard half of the blade. The blade is clamped at the root.
The maximum design voltage for the AMR blade is 3000Vpp, 0VDC.
was chosen based on depoling properties of the AFCs.
This voltage level
The AFCs had been shown to depole
between -1500V and -2000V, and therefore the AFCs could not be driven below -1500V.
To
investigate the depoling voltage of the AFC packs in the blade, the process blade is pushed to
4000Vpp, 0VDC.
at BS30.0.
Actuation data is collected using the internal torsional strain gage bridge
This bridge is located under the active area.
The conditioning amplifiers (as
described in Section 4.7) have a gain of 1000. The bridge is calibrated versus the data obtained
with the laser displacement sensors.
1.2480 /m/V.
hysteresis curve.
The calibration factor of the torsion bridge at BS30.0 is
Depoling of the AFCs is manifested as a loop at one end of the actuation
Figure 5-21 shows that he depoling voltage of the AFCs in the blade is
approximately -1500V.
The 3000Vpp, OVDC plot shows very little depoling occurring, while
the 3600Vpp and 4000Vpp curves show significant depoling loops.
Higher driving voltages than 3000Vpp can be used by adding a DC offset to the drive signal.
Such an offset was not considered in the blade design because of the drive electronics that will
be used in the wind tunnel test.
An offset requires two amplifiers per blade while only one
amplifier is available for each blade.
With only one amplifier per blade, the leadouts on the
-450 packs are simply flipped to create an electric field in the pack that is 1800 out of phase from
the +45' pack drive signal.
Adding an offset on this single amplifier would create a positive
offset in the electric field in the +450 packs and a negative offset in the -45' packs. This would
push the electric field in the -45' packs even lower, leading to depoling.
The testing setup for the process blade allows the use of two amplifiers, and therefore allowed
a DC offset to be applied to the drive signal. Applying even a small DC offset to the drive signal
can significantly reduce the depoling of the AFCs, even when driven at 4000Vpp.
119
This can
3
o
2.5 I.
Twist prediction, no correction factor,
extrapolated to 12 packs
Twist prediction, 10% reduced actuation,
extrapolated to 12 packs
Twist prediction, 25% reduced actuation,
extrapolated to 12 packs
2
1.5
Measured twist, 11 of 12 packs operational
* Measured twist, extrapolated to 12 packs
F
CO
1
0.5 6
1Hz Actuation
0
0
500
1000
1500
2000
2500
3000
Voltage (Vpp)
Figure 5-18: Actuation of the process blade at 1Hz.
data measured at BS45.5".
11/12 packs operational.
Displacement
3.5
0
3 pRev (68.4Hz)
A
2 pRev (45.6Hz)
0Q
1 pRev (22.8Hz)
--
1Hz
3
E
2.5
1--%
Cr
CO)
2
-
1.5
1
0.51
0
0
500
1000
1500
2000
2500
3000
Voltage (Vpp
Figure 5-19: Actuation data at several frequencies. Data taken with only 8 of 12 AFC packs
operational. Displacement data measured at BS45.5".
120
0.71
0.6
1Hz
2000Vpp, OVDC,
0.5
.80)4-o
~0'0
0
0.2
0.1
Active Region
20
2
3
Root
Blade station (in)
Figure 5-20: AMR blade twist versus span data. The twist is seen varying linearly through
the active region. Vibrometer data at right shows the twist distribution in the blade.
be seen clearly in Figure 5-22, where a 200VDC offset greatly reduced the size of the depoling
loop. A 500VDC offset eliminates the loop entirely.
5.3.3
Blade Mechanical Properties
The bending and torsional stiffnesses of the active area are found using the stiffness rig shown
in Figure 5-23.
A pulley-weight system applies a torsional moment on the blade for torsional
stiffness measurement.
The laser displacement sensors are used to calculate the twist rate of
the blade along the span. See Section 4.7.
The pulley system can be configured to apply a
bending moment to the blade to determine the bending stiffness of the blade.
The torsional stiffness is found by determining the twist rate of the blade as a result of an
applied load.
BS54".
For this testing, a torque of 3.37Nm is applied to the blade at blade station
The twist rate of the blade is determined by measuring the twist of the blade at
different blade stations due to the applied moment.
The torsional stiffness, GJ, is given by,
T
GJ = ,
where T is the applied torque and 0' is the twist rate of the blade [Crandall, 1976].
121
(5.1)
1.2
1.2
1.4
1
1.2
3000Vpp, OVDC, 1PerRev
0.8
3600Vpp, OVDC, 1PerRev
0.8
0.6
0.6-
0.4
0.4
C 0.2
0.2 -
0
0 -
4000Vpp, OVDC, 1 PerRev
1
0.8
0.6
0.4
Depoling loop
0.2
-0.2 -
-0.2
-0.4
-0.4 -
-0.4
-2
-1.5
-1
-0.5
0
0.5
1
1.5
2
-2
Depoling
loop
0
-0.2
-1.5
-1
-0.5
0
0.5
1.5
1
-2
2
-1.5
-1
Voltage (V)
Voltage (V)
0.5
-0.5 0
Voltage (V)
1
1.5
2
Figure 5-21: Depoling loop is seen forming as the AFC exceeds -1500V. To prevent long
term permanent damage to the AFC, the AFC should not be subjected to voltages lower than
-1500V.
1.4
1.4-
1.2
1.2
1
4000Vpp, OVDC, 1PerRev
-pp
4000Vpp, 200VDC, 1 PerRev
OVC Pr
0.8
0.8
0.6
0.6.
0.4
0.4
0.2
0.2
0
0
-0.2
-0.2
-0.4
-0.4
-0.61
-2
-e%
-2
-1.5
-1
-0.5
0
0.5
Voltage (V)
1
1.5
2
.
,
1.5
.
4000Vpp,
1
.
.
500VDC,
1 PerRev
-
0.5
0
-0.5
-1.5 -1
-0.5
0 0.5 1
Voltage (V)
1.5
2
2.5
-1
-1.5
-1
-0.5
0
0.5 1
Voltage (V)
1.5
2
2.5
Figure 5-22: Effect of a DC offset on the depoling behavior of the blade AFCs. Even small
offsets can drastically reduce the depoling of the AFCs.
The bending stiffness is found by applying a 19.5N load at blade station BS43.5", and
measuring the displacement through the active span. The bending stiffness at a given spanwise
location, EI(x), would normally be determined by finding the curvature of the blade, =,
and
using the relation,
EI(X)g
d2=
M(x),
(5.2)
where M is the moment at a spanwise location x, and u is the displacement at measured at
x.
However, since finding the curvature of the blade requires taking two derivatives of the
displacement, noise in the displacement data is greatly amplified.
122
This makes it difficult to
......
.....
Figure 5-23: Torsional stiffness testing setup. A pulley system applies a moment to the blade.
find a reliable value for the bending stiffness.
However, finding the average bending stiffness
of the blade does not require taking any derivatives of the displacement data, and can provide
better bending stiffness data. The average bending stiffness of the blade, EI, is found as,
El =
-(-X3
6u
+ 3x 2 a),
(5.3)
where P is the applied load, a is the spanwise location (from the root clamp) at which the load
is applied, x is the location on the blade at which the displacement measurements are taken,
and u is the displacement at location x [Crandall, 1976].
The value of EI found based on
displacement measurements at a given x constitutes the average bending stiffness of the blade
between the root clamp and x.
The results of the stiffness testing are:
" GJ = 162 Nm 2, 0.34R to 0.65R,
" EI
=
172 Nm 2.
The predicted values for the torsional stiffness and the bending stiffness are,
123
00
-45 packs (6 of 6), 2PRev
+45 packs (5 of 6), 2PRev
0.06
0.05
0.05
~?0.04-
0 0.04
=
0.030.030.02
0.02
-
~
0.01-
O
0.01-
00
1
1
2000
20
30
3500
00
Voltage (Vpp)
50
1000
1500
2000
2500
3000
3500
Voltage (Vpp)
Figure 5-24: Current draw of +45 and -45 plies. +45 ply has greater current draw despite
having fewer packs. The ideal capacitory line represents the current draw expected based on
low-field capacitance of the plies.
" GJ
=
155 Nm 2 ,
" El = 190 Nm 2 [Weems, 1998].
Good model-experiment correlation was achieved.
Both the torsional stiffness and the
bending stiffness measurements were within 10% of the predicted values.
Correlation with
models indicates that the ply properties that are used are accurate. A plot of bending stiffness
versus span, as defined by Eq. 5.3, is provided in Appendix C, Figure C-1.
5.3.4
Blade Current Draw and Heating
AFC testing, as described in Chapter 3, showed that the 10-mil fiber AFCs draw more current
than predicted by a capacitor model, and are subject to severe heating.
severe enough to cause the AFCs to fail.
The heating can be
It is necessary to determine if these problems are
present in the blades. High current draw could limit the upper operating envelope of the blades,
and excessive heating of the blade structure could lead to AFC failures and to degradation in
the passive material.
Figure 5-24 shows the current draw of the +45 and -45 plies in AMR blade.
includes only 5 packs, as one pack failed at 3000Vpp.
124
The +45 ply
The -45 ply has a low-field capacitance
of 47.3nF and the
+45 ply, 37.9nF. The +45 ply has higher current draw than the -45 ply,
despite having fewer packs and lower low-field capacitance. The average current draw per AFC
pack at 3000Vpp, 2PRev is 11.6mA. When extrapolated from 2PRev (45.6Hz) to 50Hz, this
current level is consistent with the average pack current shown in Figure 3-20. The performance
envelope given limited current supply shown in Figure 3-20 is therefore valid.
In the blade, removing heat from the outer layer of AFCs should not pose problems since
it is very close to the surface, but the inner ply may not be able to give off enough heat
to remain cool.
To monitor the heating of the blade, a thermistor 9 was embedded between
the spar skin and the spar core.
in temperature.
The thermistor changes resistance in response to changes
The change in resistance is significant enough that the resistance can be
monitored using standard resistance measuring devices, such as a digital multi-meter, and a
bridge circuit does not have to be built. The thermistor is located at 40% span, or BS24.0, on
the lower surface of the blade.
Initial heating tests were allowed to run indefinitely. However, during testing at 3000Vpp,
2 PRev, the blade reached 180'F and two actuator packs failed. Failures at approximately the
same temperature had been observed in unlaminated AFC testing, indicating that a component
of the AFCs may be breaking down or changing properties at high temperature.
It was decided
thereafter that the blade should not be allowed to run above 120*F.
Figure 5-25 shows heating of the blade at 2000Vpp, OVDC at different frequencies.
heating tends to grow relatively linearly with frequency.
drive voltage.
The
Figure 5-26 shows heating data versus
The heating increases drastically as 3000Vpp is approached.
The increase in
blade temperature is much larger between 2600Vpp and 3000Vpp than between 2000Vpp and
2600Vpp.
Figure 5-27 shows that blade heating can be reduced significantly by applying a DC offset
to the drive voltage. A 500VDC offset on a 2600Vpp signal can reduce the blade temperature
increase during operation by approximately 30%.
Depoling of domains in the AFC is one of the major sources of current draw of the AFCs.
As the domains pole and depole, the shift charge from one portion of the domain to another,
resulting in a net current. A DC offset decreases the amount of depoling that occurs in the
9 ETG-50A, Micro-Measurements Group, Inc., Raleigh, North Carolina.
125
125
-
120
5PRev
115 -
3PRev
U_
6110
2PRev
-
(105
0.100 E
1 PRev
95 -
80 -
0
2
4
6
8
10
12
14
16
18
20
Time (min)
Figure 5-25: Blade heating as a function of drive frequency.
mately linearly with frequency.
The heating increases approxi-
130.
125-
120-
3000Vpp
6115110-
0
W105-
0.
2600Vpp
E
010-
951-
2000Vpp
90-
OVDC, 1PRev
8
5
Time (min)
10
1
Figure 5-26: Heating of the AMR blade at different drive voltages.
126
110
OVDC
105 -
20OVDC
LL'
1
00
-
50OVDC
aE 950
o)
2600Vpp, 1PRev
0
6
4
2
14
12
10
8
Time (min)
Figure 5-27: Applying a DC offset to the drive voltage, the blade heating can be reduced.
0.04
,
1
1
1
0.0
,
1
1
60OVDC
0.03
0.04
OVIDC
0.02
20VDC
0.03
.2OVDC
--
0.020
0.01
-0.01
-0
~
0
000D
-0.01
20OVD
70VVC
I
-0.01
OVD
-001
500VDC
-0.02
-0.03 -
-0.03
4000Vpp, 1PRev
2600Vpp, 2PRev
-1.5
-1
-0.5
0
0.5
1
1.5
2
-2
-1.5
-1
-0.5
0
0.5
1
Voltage (KV)
Voltage (KV)
Figure 5-28: Effect of DC offset on current draw.
127
1.5
2
2.5
3
1.6 X1F
0.4
-45 plies, capacitance
1.5
-45 plies, tan6
1.4
iz
0)
0.35
7
I-
1.3Cu
0.3
a)
V
C
Cu
1.2-
7A
C
0
A
-
0.25
0.9
0.8
0.2 -
0.7
0
0
0.61
1000
500
1500
2000
2500
3000
100
3500
C
0
/7
7
200
300
4
500
600
700
800
8
00
1000 1
0.4
+45 plies, capacitance
+45 plies, tan6
7-
0.35
12
0.3 -
11
a)
CL
CO
0
900
Applied field (Vrms/mm)
Voltage (Vpp)
1
7
,7
7
7-
Cu
a) 0.25 V
C
Cu
0.2 -
109-
8-
70-
7
-
7-
0.15 -
7,
0.1 -
500
1000
150
2000
25
3000
0.05
3500
Voltage (Vpp)
-a-o
-0
-a
0
0
200
300
400
500
600
700
800
900
1000
1100
Applied field (Vrms/mm)
Projected, high free strain AFC
Projected, low free strain AFC
Blade Data
Figure 5-29: Capacitance and tan6 data for the +450 plies (5 of 6 packs operational) and -45'
plies (6 of 6 packs operational). The levels expected based on individual laminated AFCs is
shown.
128
AFC by decreasing the fields opposing the polarization of the piezoelectric material. Therefore,
increasing the minimum voltage during the drive cycle can significantly decrease the current
draw of the AFC. Figure 5-28 shows two graphs: 2600Vpp with variable offset, and 4000Vpp
with variable offset.
creased.
As the current curves shift right, the maximum current draws are de-
The very severe effects of depoling can be seen in the 4000Vpp graph.
The current
is very large when no DC offset is used.
The larger current corresponds to the depoling loop
at negative voltage seen in Figure 5-22.
When even a small DC offset, 200V, is applied to the
drive signal, the current is drastically decreased.
This lower current draw occurs because the
AFCs are not driven negative enough to cause bulk depoling.
Capacitance and dielectric loss factor data is found for the actuators in the blade using
the procedure presented in Chapter 3. Figure 5-29 shows the blade capacitance and tan6 data
along with the expected levels based on individual laminated AFC data. Data from both high
free strain and low free strain AFCs is presented in Figure 5-29.
In Chapter 3, high free strain
packs were shown to have greater dielectric constant and dielectric loss factor than low free
strain packs. The blade data is expected to fall between these two expected levels. The figure
shows that the +45'
laminated AFCs.
AFC testing.
plies conform to expectations based on data collected from individual
The -45' plies show slightly lower capacitance and tan6 than expected from
Based on average free actuation levels for the -45* and +45' plies, 1267wE and
1424pE respectively, it is expected that the
+45* plies should show a higher capacitance and
dielectric loss. The experimental data follows that expectation.
5.3.5
Actuator Pack Failure History
No actuators failed during testing through 2000Vpp.
3000Vpp, actuator pack failures began occurring.
When the voltage was increased to
Table 5.1 shows a summary of the pack
failures in the blade. Figure 5-30 shows the numbering scheme of the active actuator packs in
the process blade, and the location of the failures.
Actuator pack 9 failed almost immediately as the blade was being driven at 3000Vpp for
the first time.
Packs 2 and 3 failed simultaneously during high temperature testing.
During
heating testing at 3000Vpp, 2PRev, the blade climbed past a temperature of 180'F. Packs
2 and 3 failed simultaneously at this temperature.
129
Pack 6 failed during subsequent heating
Pack Number
9
2
3
Location
Lower surface, -45' ply
Upper surface, 450 ply
Upper surface, 45' ply
Description
Failed after -1 sec. at 3000Vpp
3000Vpp, 2Prev, 180'F blade temperature
3000Vpp, 2Prev, 180'F blade temperature
6
5
Upper surface, -45' ply
Upper surface, 450 ply
2000Vpp, 3Prev, 120'F blade temperature
4000Vpp, 1000VDC, after ~5 min.
Table 5.1: Failure history of process blade actuator packs.
Upper 45, Outer ply
Upper -46, Inner ply
Lower 45, Inner ply
/
10
12
Lower -46, Outer ply
7
Failed actuator pack
/1117
Live actuator pack
Figure 5-30: Schematic of actuator pack failures.
testing, at about 120'F while running at 2000Vpp, 3Prev.
After all data had been collected on the process blade, the drive voltage was pushed higher
than 3000Vpp.
As the voltage was increased, a DC offset was applied to the drive voltage to
prevent depoling.
The blade was run at 4000Vpp, 1000VDC at 1PRev for 5 minutes before a
failure occurred.
During this test, a fan was used to blow air over the blade to provide some
cooling. The temperature measured by the embedded thermistor did not exceed 100'F.
Possible causes for the actuator pack failures include AFC manufacturing defects, blade
manufacturing damage, and decomposition of AFC materials at high temperatures.
Small defects in the AFCs could lead to low resistance paths between electrode fingers.
Small voids could create an arcing path between electrode fingers. This type of defect may not
pose any danger of failure at low voltages, but as the electric field strength is increased, the
failure becomes more likely to occur.
130
Stresses during blade manufacturing are very large. Shear forces could cause delaminations
in the packs, and could lead to electrical failure.
Also, high curvature of the airfoil on the
upper surface of the blade may cause fiber cracking in the AFCs. This cracking could damage
the electrode pattern.
Any thinning or cracking of the electrode pattern could lead to a local
heating of the electrode due to current flow through a damaged conductor.
As the current
draw of the AFC increases, the local heating will become more significant, and will eventually
lead to breakdown.
The thermistor in the blade is located on the lower surface of the blade, under the midspan
of the active area. It is possible that heating can occur in an area removed from the thermistor.
This heating would not be observable on the thermistor measurement.
It is therefore possible
that temperatures that are much higher than the thermistor measurement could occur. These
high temperatures could lead to localized degradation of the AFC matrix, and may cause
breakdown.
Actuator packs 2, 3, and 6 may have failed because of damage caused by high temperature in
the blade. It is likely that the 180'F temperature observed on the thermistor is not responsible
for any degradation of the AFC properties, but that localized higher temperatures may have
been to blame.
Actuator pack 9 probably failed because of manufacturing defects.
A small void may have
been present in the pack, or a local thinning in an electrode finger may have been to blame.
This pack did not fail because of large scale heating, as it happened after running for only about
a second at 3000Vpp, 1PRev.
Actuator pack 5 was exposed to very high voltages.
Running the blade with a 1000VDC
offset, at 4000Vpp, meant that the highest voltage seen between the electrode fingers was 3000V.
This is double the maximum voltage that the actuator packs will be exposed to during normal
testing.
The high voltage cycle was meant to push the performance to failure to determine
what margin of safety is present in driving the active material at the design voltage of 3000Vpp.
It should be noted that the failures to the AFCs were very difficult to detect at low voltages.
Historically, failures have been detected by measuring resistance and capacitance of each pack.
Failures generally appeared as low resistances across a pack. However, when failures occurred in
the blade, the failed packs often appeared normal when measuring capacitance and resistance.
131
The failed packs could only be detected by running each pack to high voltage individually. The
failed pack would fail at 50% to 75% of the voltage at which the first failure occurred.
This
method of failure detection is undesirable because it could potentially further damage the blade.
Any damage that occurred to the AFC and the blade as a result of the first failure would be
aggravated. However, there is no alternative to the damage detection procedure. It is intrinsic
to the 10-mil fiber AFCs.
5.3.6
Destructive Testing
Upon completion of all mechanical and electrical testing, the blade was cut into sections to
allow inspection of the internal blade features.
The AFCs showed no out of plane deformation.
Two areas were of particular interest:
the overlap over the leading edge weight and the area where the fairing skin overlaps the spar.
Each location could cause deformation of the AFC due to steps in the laminate.
Figure 5-31
shows a typical cross-section of the blade. The AFC can be seen in the spar, with no significant
out of plane deformation.
The only damage that was observed in the AFCs was at the location of the failure of actuator
pack 5.
This failure had made a visible bubble on the surface of the blade, and upon cross-
sectional inspection, it was observed that the AFC had delaminated. This could cause concern
in flight, but since the passive plies were intact, it should not be catastrophic.
pack 5 was extensive, probably due to the very high driving voltages.
The damage to
Of the other packs that
failed, two made small raised areas in the spar skin, and two did not show any physical signs
of damage.
Cuts were made through several electrical connections.
The copper strips have a stress
relief in them that is a natural outcome of the manufacturing procedure.
The copper strips
are initially rather taut when the connections are made, but as the fairing is compressed, the
strips develop a stress relief that prevents breaking of the strip.
Figure 5-33 shows the stress
relief clearly.
One area of concern is the transition of the flexible circuit from the web to the fairing. As
the fairing was compressed, some of the transitions became rather sharp. However this should
not pose major problems.
The flexible circuits are designed to be bent sharply, as long as the
132
Figure 5-31: Typical crossection of the blade.
Figure 5-32: Crossection through failure of pack number 5.
Figure 5-33: Crossection of connections.
the flexible circuit.
Copper strips can be seen running from AFC tab to
Figure 5-34: Crossection of flex transition area. The radius has been distorted on the bottom
surface, and has been pinched at the upper surfa
133
bend is only made once [Allflex, 1998].
Also, the adhesive used on the web of the blade and
under the flexible circuits holds the circuit in place, and doesn't allow it to bend repeatedly.
See Figure 5-34. Foaming adhesive may be used underneath the flexible circuit radius to push
the radius outward and now allowing it to collapse.
5.3.7
Windtunnel Testing Recommendations
Prior to hover testing, the blades should be bench tested in the manner described in this chapter.
Actuation testing should be performed at increasing voltages to determine the non-rotating
twist rates of the blades.
This testing will serve both to provide data for model-experiment
correlation and to insure that the blades survive the range of voltages that will be seen in
testing.
During windtunnel testing of the AMR blade, blade temperature and current should be
monitored closely.
If the blade temperature rises significantly, testing should be stopped,
or the drive level should be decreased.
It is likely that heating will not be as significant
during hover testing as it was in bench testing due to the airflow over the blades.
Given
that enough current is available and that the blade temperature does not approach 180'F, few
problems should be encountered with the AFCs in the blade at 3000Vpp and below.
It has
been observed that pack failures generally occur within the first few seconds of reaching a given
voltage.
Therefore, if a pack doesn't fail within the first few seconds at 3000Vpp, generally it
will not fail at that voltage. If damage occurs to the pack, whether mechanical or chemical or
thermal, damage may occur after longer periods of testing.
Current should be monitored to ensure that no sharp increases in current draw occur during
testing. Such an increase could indicate that a low resistance path is developing across an AFC,
and is therefore indicative of an impending pack failure.
It is likely that such damage to an
AFC is irreversible, but it may be desirable in certain circumstances to shut the rotor down
before an AFC failure occurs.
Amplifiers with a current trip should be used during testing.
If a failure occurs and the
current level rises above what is expected during normal operation, the amplifier should be
shut down immediately and automatically.
Using a current limit feature, in which the peak
current is limited, does not sufficiently protect the AFCs.
134
If such a feature is used, current
will continue running across a failure until the amplifier is shut off manually. This could cause
catastrophic damage to the blades.
The failure will tend to burn as current runs across it,
and if the current is not shut down within a few milliseconds, the burning could spread to the
passive plies in the blade and compromise the structural integrity of the blade.
135
136
Chapter 6
Conclusions and Recommendations
6.1
Summary and Conclusions
This thesis has presented developments made in the understanding, design, and manufacturing
of integrally actuated composite rotor blades.
actuation technology were developed:
Two main areas of integral helicopter blade
actuator understanding, including power consumption
and the effects of clamping on actuation, and blade manufacturing, focusing on the power bus
design to improve the manufacturability of the blades.
The first portion of the thesis focused on Active Fiber Composite actuator testing.
In
the first generation MIT/Boeing integrally actuated blade, it was observed that the actuation
authority consistently fell short of expectations.
However, no studies had been done to un-
derstand the behavior of the material when laminated.
Such studies were conducted as part
of this work, and indicated that the actuators performance is non-linear with voltage and constraint level. It was found that to correct linear blade models for the non-linearity of the AFC
actuators, reduction factors had to be applied to the actuation capability of the actuators in the
models. For the AMR blade AFCs, these reduction factors range from about 10%, for low free
strain actuators, to 25%, for high free strain actuators.
Failure to use such reduction factors in
modeling may lead to significant overpredictions of actuation capability. The exact reduction
factors necessary for accurately modeling AFCs can only be determined experimentally.
As a
result, several AFCs must be tested from each batch of AFCs to determine the cutdown factors
necessary in modeling.
137
Current draw and heating studies were performed on the AMR blade AFCs.
Historically,
first order approximations of current draw have been calculated using an ideal capacitor model,
and using the low field capacitance of the AFCs. It was found, however, that the current draw
of the AFCs can be as much as 3 times greater than predicted by the ideal capacitor model.
By fitting a parallel capacitor-resistor circuit to the experimentally determined current draw,
it was observed that the capacitance of the AFCs increases by as much as 100% from low field
(500Vpp) to high field (3000Vpp).
It was also observed that the dielectric loss factor of the
AFCs increases from approximately 0.15 at low fields (500Vpp) to approximately 0.35 at high
fields (3000Vpp).
Along with the increased current draw and greater loss at high fields comes
significant self heating of the AFCs. Addition of DC offsets to the driving fields can decrease
the current draw and heating of the AFCs.
The second part of the thesis focused on the changes made to the manufacturing process and
to the blade design to improve reliability of the blades. The first generation integral blade was
handicapped by many actuator and power bus failures. Many of these failures were attributed
to poor connectivity to the power distribution circuit. By improving the manufacturing process
and AFC pack design to enable better connections to be made in the blade, the survival rate of
the actuator packs was drastically improved, and ease of blade manufacturing was increased.
By moving the power lines, in the form of a flexible circuit, from the narrow web of the blade
to the surface of the fairing, access to the connections was improved drastically and the need
for a sharp bends in the electrode tab was eliminated.
The third part of the thesis presented the bench testing of the first blade built with the new
power bus design and with the new 10-mil fiber AFCs.
The testing was a success, with all
connections and AFCs surviving the cure and operating at 2000Vpp.
One actuator pack was
lost at 3000Vpp. Such failures may be inevitable. Since the driving voltages are so high, small
manufacturing defects, or slight damage to the packs during the cure can cause these failures
to occur.
Blade current draw was in line with expectations based on individual pack testing. Blade
heating was significant, but also in line with expectations.
The internal temperature of the
blade climbed quickly at high voltage and frequency, and was the cause of two, possibly three,
AFC failures in the blade.
These failures may have been due to local temperatures that were
138
much higher than the blade temperature, and my therefore not have been detectable with
the internal temperature measuring device.
To prevent such failures from occurring,
blade
temperature should be monitored closely and not allowed to reach more than 120'F. Heating
should not prove to be problematic during hover and windtunnel testing of the blades. The air
flow over the blades should keep the blades cool. However, care should be taken during bench
testing and non-rotating blade testing to not exceed a blade temperature of 120'F.
Blade actuation closely matched expectations.
The blade achieved a twist rate of 2.3'/m
at 3000Vpp, 1Hz, while the model prediction was 2.54/m, less than 10% higher. The actuation
modeling was done with a 10% reduced actuator free strains to account for the high-field nonlinearities observed in the AFC testing.
6.2
Recommendations for Future Work
Further work should be done to improve the manufacturing procedure of the blades. Although
improved over the first generation blade, the second generation power bus and connection design
is still prone to damage during the manufacturing process. The electrode tabs protruding from
the spar during the blade cure could become damaged when the spar is removed from the mold
following the spar cure, or during the fairing cure. Eliminating the electrode tab protruding
from the packs, leaving only two copper strips exiting the pack could reduce the risk of breakage
during manufacturing.
The actuator pack design and manufacturing procedure would have
to be changed to allow embedding of the copper strip without having the electrode pattern
protruding from the pack.
Having only the strong and malleable copper strips exiting the
packs would eliminate the need for the depressions made in the fairing foam to accommodate
the electrode tabs, simplifying the manufacturing procedure.
The study conducted in Chapter 3 on the non-linearity of mechanical and electrical behavior
AFCs suggests that changing the fiber ceramic from PZT-5A to a harder ceramic, like PZT-4,
may be beneficial and should be investigated.
The mechanical and electrical properties of
PZT-4 makes it more suited for high power applications than PZT-5A.
constant and dielectric loss factors are lower than those of PZT-5A.
First, the dielectric
Moving from PZT-5A
to PZT-4 would therefore reduce the current draw and self heating of the AFCs.
139
A lower
current draw would allow the performance envelope to be expanded to higher voltages and
higher frequencies given a limited current availability.
Less heating of the AFCs will help
prevent structural degradation (resin softening) that occurs at high temperatures. A drawback
of using PZT-4 fibers is that they have approximately 50% lower free strain than PZT-5A.
However, the material behavior is more linear in the voltage range in which the blades are
driven, and is significantly less hysteretic.
The higher stiffness and lower non-linearity will
help offset some of the lower actuation capability of PZT-4 compared to PZT-5A. A collection
of data on PZT-4 and PZT-5A is included in Appendix E for reference.
The corrections factors necessary in modeling constrained AFCs should be studied.
thesis has quantified the reduction factors.
This
However, the mechanism causing the non-linearity
in the actuation authority under free and laminated conditions is not understood.
Specifically,
the difference in the reduction factors needed in modeling high free strain and low free strain
AFCs should be investigated. A better understanding of these non-linearities can lead to better
modeling capability of embedded actuators.
The continuation of the work presented in this thesis will involve hover and forward flight
testing of the AMR blade.
conducted.
Blade characterization and vibration reduction studies will be
Mechanical and electrical fatigue testing will also be performed on the blades to
determine how the blade will perform through long term operation.
140
Bibliography
[Allflex, 1998]
Design Guide, AlIflex, inc. Northfield, MN, 1998.
[Allflex, 1999]
Allflex, inc., personal correspondence, 1999.
[Amer, 1974]
Amer, K.B., Neff, J.R., "Vertical-plane pendulum absorbers for minimizing helicopter vibratory loads," Proceedings of Rotorcraft Dynamics Conference, Moffett Field, CA, NASA Ames Reserach Cen-
ter,, Feb. 1974, pp. 219-222. N74-34510.
[Bent, 1997]
Bent, A.A., Hagood, N.W., "Active Fiber Composites for Structural
Actuation" AMSL #97-1,
Massachusetts Institute of Technology,
January, 1997.
[Berlincourt, 2000]
Berlincourt,
erties
of
D.,
Krueger,
Morgan
Publication
226,
H.H.,
Matroc
Near,
Piezoelectric
C.,
"Important
Ceramics,"
Morgan Electro Cereramics,
www.morganelectroceramics.com/techpubl.html,
[Berlincourt, 2000b]
Berlincourt,
tric
D.,
Ceramics
Conditions
lication
Krueger,
under
Various
of Radiating
228,
Morgan
H.H.,
Sonar Transducers,"
Electro
Cereramics,
Technical
Bedford,
Ohio,
2000.
"Behaviour
Environmental
Prop-
of
Piezoelec-
and
Operation
Technical
Bedford,
PubOhio,
www.morganelectroceramics.com/techpubl.html, 2000.
[Bielawa, 1992]
Bielawa, R.L., Rotary Wing StructuralDynamics and Aeroelasticity,
AIAA Education Series, Ed. Przemieniecki, J.S., 1992.
141
[Braithwaite, 1990]
Braithwaite, N., Weaver, G., Electronic Materials,Alden Press, Lon-
don, 1990.
[Braun, 1984]
Braun, D., "Ground and flight tests of a passive rotor isolation system for helicopter vibration reduction," Vertica, Vol. 8, No. 1, 1984,
pp. 1-14.
[Bussom, 2000]
Bussom, R.C., Boeing Helicopters, Philadelphia, PA, personal correspondence, 2000.
[Centolanza, 2000]
Centolanza, L.R., Smith, E.C., "Design and Experimental Testing of an Induced-Shear Piezoelectric Actuator for Rotor Blade
Trailing Edge Flaps." AIAA-2000-1713.
AIAA/ASME/ASCE/AHS/ASC
Proceedings of the 41st
Structures, Structural Dynamics
and Materials Conference, April 3-6, 2000. Atlanta, GA.
[Cesnik, 1999]
Cesnik, C.E.S., Morales, M.O., "Active Composite Beam CrossSectional Modeling - Stiffness and Active Force Constants," AIAA-
1999-1548, Proceedings of the 41st AIAA/ASME/ASCE/AHS/ASC
Structures, Structural Dynamics and Materials Conference, 1999.
[Cesnik, 1998
Cesnik, C.E., Shin, S., "On the Twist Performance of a Multiple-Cell
Active Helicopter Blade," Proceedings of the 9th InternationalConference on Adaptive Structures and Technologies, Cambridge, MA,
14-16 October 1998.
[Chen, 1994]
Chen, P.C., Chopra, I., "Induced Strain Actuation of Composite
Beams and Rotor Blades with Embedded Piezoceramic Elements,"
SPIE Proceedings of the Smart Structures adn Intelligent Systems
Conference, Vol. 2190, 1994, pp. 123-140.
[Crandall, 1976]
Crandall, S.H., Dahl, N.C., Lardner, T.J., An Introduction to the
Mechanics of Solids, McGraw-Hill Book Company, New York, 1976.
142
[Derham, 1996]
Derham, R.C., Hagood, N.W., "Rotor design using smart materials
to actively twist blades," Proceedings of the 52nd Annual Forum of
the American Helicopter Society, 1996.
[Derham, 2000]
Derham, R.C., Boeing Helicopters, Philadelphia, PA, personal correspondence, 2000.
[Desjardins, 1978]
Desjardins, R.A., Hooper, W.E., "Helicopter rotor vibration isolation," Vertica, Vol. 2, No. 2, 1978, pp. 145-159.
[DuPlessis, 1995]
DuPlessis, A.J., Hagood, N.W., "Performance Investigation of Twist
Actuated Single Cell Composite Beams for Helicopter Blade Control", Proceedings of the 6th International Conference on Adaptive
Structures Technology, Key West, FL, 1995.
[Friedmann, 1998]
Friedmann, P. P., "Rotary-Wing Aeroelastic Scaling and Its Applications to Adaptive Materials Based Actuation", Proceedings of
the 39th AIAA/ASME/ASCE/AHS/ASC
Structures, StructuralDy-
namics and Materials Conference, April 20-23, 1998. Long Beach,
CA.
[Ghandi, 1998]
Ghandi, K., Hagood, N.W., "Non-linear Modeling and Characterization Techniques for Phase Transitions in Electromechanically Coupled Devices" AMSL #98-1,
Massachusetts Institute of Technology,
May, 1998.
[Giurgiutiu, 2000]
Giurgiutiu, V., "Recent Advances in Smart-Material Rotor Control Actuation.",
AIAA-2000-1709.
AIAA/ASME/ASCE/AHS/ASC
Proceedings of the
41st
Structures, Structural Dynamics
and Materials Conference, April 3-6, 2000. Atlanta, GA.
[Glenn, 2001]
Glenn, T.S., Ph.D. thesis in preparation, Department of Aeronautics
and Astronautics, Massachusetts Institute of Technology, 2001.
143
[Ham, 1987]
Ham, N.D., "Helicopter individual-blade-control research at MIT
1977-1985," Vertica, Vol. 11, No. 1/2, 1987, pp. 109-122.
[Harper, 1999]
Harper, J.E., Hagood, N.W., "Analysis of Non-linear Electroelastic Continua with Electronic Conduction," AMSL #99-2,
Massa-
chusetts Institute of Technology, May, 1999.
[Jacklin, 1995
Jacklin, S.A., Blass, A., Teves, D., Kube, R., "Reduction of Helicopter BVI Noise, Vibration, and Power Consumption Through
Individual Blade Control," Proceedings of the 51st Annual Forum of
the American Helicopter Society, Vol. 1, 1995, pp. 662-680.
[Jaffe, 1971]
Jaffe, B., Cook, W. R., Jaffe, H., Piezoelectric Ceramics, Academic
Press, 1971.
[Jones, 1999]
Jones, R.M, Mechanics of Composite Materials, Taylor & Francis,
1999.
[Pizzocchero, 1999]
Pizzocchero, A., Continuum Control Corporation, Billerica, MA,
personal correspondence, 1999.
[Prahlad, 2000]
Prahlad, H., Chopra, I., "Development of an Adaptive Flexbeam
for Rotorcraft
Alloy
Applications
Actuators."
Using
AIAA-2000-1712.
AIAA/ASME/ASCE/AHS/ASC
Embedded
Shape
Proceedings of
Memory
the
41st
Structures, Structural Dynamics
and Materials Conference, April 3-6, 2000. Atlanta, GA.
[Prechtl, 2000]
Prechtl, E.F., Hall, S.R., "Design and Implementationof a piezoelectric Servo-Flap Actuation System for Helicopter Rotor Individual
Blade Control," AMSL #00-3, Massachusetts Institute of Technology, January, 2000.
[Prechtl, 2000b]
Prechtl, E.F., Hall, S.R., "Closed-Loop Vibration Control Experiments on a Rotor with Blade Mounted Actuation." AIAA-2000-1714.
Proceedings of the 41st AIAA/ASME/ASCE/AHS/ASC
144
Structures,
Structural Dynamics and Materials Conference, April 3-6, 2000. Atlanta, GA.
[Reichert, 1980]
Reichert, G., "Helicopter Vibration Control - A Survey," Proceedings
of the Sixth European Rotorcraft and Powered Lift Aircraft Forum,
Bristol, England, 16-19 September, 1980.
[Rodgers, 1998]
Rodgers, J. P., Hagood, N.W., "Development of an Integral Twist
Actuated Rotor Blade for Individual Blade Control." AMSL #98-6,
Massachusetts Institute of Technology, October 1998.
[Roglin, 1994]
Roglin, R.L., Hanagud, S.V., Kondor, S., "Adaptive airfoils for helicopters," Proceedings of the AIAA Adaptive Structures Forum, 1994.
AIAA Pater 94-1764.
[Schmidt, 2000]
Schmidt, M.C., Hagood, N.W., "Design and Manufacturing of a
Second Generation Integral Twist-Actuated Rotor Blade", Proceed-
ings of the AIAA-2000-1710, 41st AIAA/ASME/ASCE/AHS/ASC
Structures, Structural Dynamics and Materials Conference, April 3-
6, 2000. Atlanta, GA.
[Sensor Technology, 1995] Sensor Technology Limited, Collingwood, Ontario, Piezoelectric Ceramics Procuct Catalogue and Application Notes, 1995.
[Shaw, 1981]
Shaw, J., Albion, N., "Active Control of the Helicopter Rotor for
Vibration Reduction," Journal of the American Helicopter Society,
Vol. 26, No. 3, 1981, pp. 32-39.
[Shin, 1999]
Shin, S.,
Cesnik, C.E.S., "Design, Manufacturing and Testing of
an Active Twist Rotor." AMSL #99-3,
Masachusetts Institute of
Technology, June 1999.
[Straub, 2000]
Straub, F.K., Kennedy, D.K., Domzalski, D.B., Hassan, A.A., Ngo,
H., Anand, V., Birchette, T., "Smart Material Actuated Rotor
145
Technology - SMART." AIAA-2000-1715. Proceedings of the 41st
AIAA/ASME/ASCE/AHS/ASC
Structures, Structural Dynamics
and Materials Conference, April 3-6, 2000. Atlanta, GA.
[Straub, 1996
Straub, F.K., Merkley, D.J., "Design of a servo-flap fotor for reduced
control loads," Smart Materials and Structures, Vol. 5, No. 1, 1996,
pp. 68-75.
[Strock, 1999]
Strock, H.B., Pascucci, M.R., Parish, M.V., Bent, A.A., Shrout,
T.R., "Active PZT fibers, a commercial production process," Proceedings of the SPIE Conference on Smart Materials Technologies,
Newport Beach, CA, March, 1999.
[Swanson, 1994]
Swanson, S.M., Jacklin, S.A., Blaas, A., Kube, R., "Individual Blade
Control Effects on Blade Vortex Interaction Noise," Proceedings of
the 50th Annual Forum of the American Helicopter Society, Vol. 1,
1994, pp. 81-101.
[Walz, 1982]
Walz, C., Chopra, I., "Design and testing of a helicopter rotor model
with smart trailing edge flaps," Proceedings of the 35th Structures,
Structural Dynamics and Materials Conference, Adaptive Structures
Forum, 1994. AIAA Paper No. 94-1767.
[Weems, 1998]
Weems, D., Boeing Helicopters, Philadelphia, PA, personal corre-
spondence, 1998-2000.
[Wickramasinghe, 2001]
Wickramasighe, V., Master of Science thesis in preparation,
De-
partment of Aeronautics and Astronautics, Massachusetts Institute
of Technology, 2001.
[Wickramasinghe, 2000b] Wickramasinghe, V., Hagood, N.W., "Material Characterization of
Active Fiber Composite Actuators for Active Twist Helicopter Rotor Blade Applications", Proceedings of the AIAA-2000-1499, 41st
AIAA/ASME/ASCE/AHS/ASC
Structures, Structural Dynamics
and Materials Conference, April 3-6, 2000. Atlanta, GA.
146
[Wickramasinghe, 2000c]
Wickramasinghe, V., Active Materials and Structures Laboratory,
MIT, Cambridge, MA, personal correspondence, 2000.
147
148
Appendix A
Coupon Testing Data
149
Head
Batchi
Batch 2
Batch 3
Sample
Number
1
Process
PreProcessing
Postprocessing
2000V/3000V
2000V/3000V
(Ae)
(AE)
Change
(%)
2
Control
Heat to 250*F
253/431
309/505
251/424
288/482
-1%/-2%
-7%/-5%
3
4
5
Laminate, 50psi
Laminate, 500psi
Laminate, 500psi
262/445
240/406
291/495
132/213
103/172
134/222
-50%/-52%
-57%/-58%
-54%/-55%
6
7
Failure
Failure
8
9
10
Control
Heat to 250'F
Laminate, 50psi
665/1125
648/1074
498/867
527/897
612/1023
134/222
-21%/-20%
-6%/-5%
-73%/-74%
11
Laminate, 500psi
476/849
117/190
-75%/-78%
12
Laminate, 500psi
520/868
144/231
-72%/-73%
13
14
Control
Heat to 250*F
628/1092
639/1098
601/1059
609/1066
-4%/-3%
-5%/-3%
15
16
Laminate, 50psi
Laminate, 500psi
664/1145
690/1186
253/408
224/359
-62%/-64%
-68%/-69%
17
Control
708/1201
653/1145
-8%/-5%
18
19
Heat to 250*F
Laminate, 50psi
639/1088
657/1109
604/1052
250/400
-5%/-3%
-62%/-64%
20
Laminate, 500psi
653/1092
209/340
-68%/-69%
Table A.1: Results of coupon testing.
150
Pack
#
Pre-Processing
Capacitance
(nF)
Post-Processing
Capacitance
(nF)
1
2
3
4
5
6
7
8
9
410
447
387
372
492
481
375
657
658
410
449
388
375
488
Failed
Failed
657
667
10
696
734
11
12
13
659
614
602
707
628
599
14
15
16
17
18
19
571
673
667
639
643
715
570
660
663
640
642
699
20
676
668
Table A.2: Capacitance data for test coupons.
151
Pack
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
#
Pre-Processing
Resistance (MQ)
Post-Processing
Resistance (MQ)
71
63
74
77
54
60
45
35
35
31
32
38
40
45
33
36
39
37
30
33
72
70
88
86
55
Failed
Failed
35
36
36
38
38
40
44
43
43
38
37
40
43
Table A.3: Resistance data for test coupons.
152
Appendix B
Blade Section 4.0 Supplement
This appedix contains mechanical drawings for the blade components of Section 4.0.
following drawings are presented:
" Section 4.0 mechanical drawing
" Section 4.0 flexible circuit drawing
" Section 4.0 tip fitting drawing
* Section 4.0 electrode pattern.
The parts list, containing ply dimensions and layup information is also presented.
153
The
Description
Spar Ply 1
Spar Ply 2AU
Spar Ply 2AL
Spar Ply 2U
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Ply 2L
Ply 2XU
Ply 2XL
Ply 2N
Ply 3
Ply 4AU
Ply 4AL
Ply 4U
Ply 4L
Ply 4XU
Py 4XL
Ply 5
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Spar
Ply 6AL
Ply 6U
Ply 6L
Ply 6XU
Ply 6XL
Ply 6N
Ply 7
Ply 8
Spar Ply 6AU
Material
E120 @ 45
S2/SP381 @+45
S2/SP381 @+45
IDEPFC @ 45
IDEPFC @ 45
S2/SP381 @+45
S2/SP381 @+45
S2/SP38i 00
S2/SP381 00
S2/SP381 @-45
S2/SP381 @-45
IDEPFC @-45
IDEPFC @ -45
S2/SP381 @-45
S2/SP381 @-45
E120 0 45
E120
Width
Length (short side)
Length (long side)
[in]
4.29
1.92
1.92
[in]
20.63
6.80
6.80
[in]
8.72
8.72
1.92
1.92
1.92
1.92
0.43
2.12
1.92
1.92
1.92
1.92
1.92
1.92
4.23
@ 45
5.80
5.80
20.63
20.63
6.30
6.30
7.72
7.72
6.30
6.30
20.63
8.22
8.22
5.80
7.72
8.22
8.22
Comments
Wraps Around Leading Edge
Upper Surface Only, Angled Outer End
Lower Surface Only, Analed Outer End
Upper Surface Only, Angled Ends
Lower Surface Only, Angled Ends
Upper Surface Only, Angled inner End
Lower Surface Only, Angled Inner End
Wraps Around Leading Edge
Wraps Around Root Pin
Upper Surface Only, Angled Oat End
Lower Surface Only, Angled Uer End
Upper Surface Only, Angled Ends
Lower Surface Only, Angled Ends
Upper Surface Only, Angled inner End
Lower Surface Only, Angled Inner End
Wraps Around Leading Edge. Cut to make
round Leading Edge
Wraps Around Leading Edge
Upper Surface Only, Angled Ends
Lower Surface Only, Angled Ends
E120 @ 45
IDEPFC @45
IDEPFC @45
E120 @45
E120 @45
S2/SP381 @0
E120 0 45
IM7/SP381 @0
1.92
1.92
1.92
1.92
1.92
0.37
4.19
2.09
5.80
7.72
6.80
6.80
20.63
20.63
20.63
8.72
8.72
Web Ply 1
Web Ply 2
Web Ply 3
El 20@45
El 20@45
E120045
1.02
1.00
0.98
20.63
20.63
20.63
Wraps around spar heel
Wraps around spar heel
Wraps around spar heel
Nose Block Filler
S2/SP381 @0
1.75
20.63
Roll around weights, form to leading edge
Lower Fairing Skin
Upper Fairing Skin
E120@4.
E120045
3.64
TE Tab
TE Stiffener
IM7/SP381 @90
IM7/SP381 @0
TE Filler
Wraps Around Leading Edge
Wraps Around Leading Edge.
Wraps Around Root Pin
20.63
20.63
Film Adhesive
0.75
0.21
0.17
20.63
Roll into rope, install at TE of fairing core
Spar Adhesive
Film Adhesive
Spar -Fairing Adhesive ilm Adhesive
4.43
1.15
20.63
20.63
Wraps around spar core
'Wraps around spar heel
S2/SP381 @0
1.00
8.00
length approximate. Will angle both
Root Doubler
Figure B-1: Parts list for Section 4.0.
ends later on.
Ply sizes, locations, and layup information is shown.
154
-(1354'
1~-Tip
1.250'
-4-0.600
Fitting
N
N
1,000'
2.050'
1.000'
0
0
4)
AFCs
20.630'
7.460'
6.960'
4600'
Li1
ITip~
L~~j
5.442'
Figure B-2: Section 4.0 drawing.
155
~It
Oq
S
Layer I
see~es JMME
C)
a A
I-I
I.A
C)
Layer
E
CD
CIT
MIT 4 Layer FlesxaLe arcut
Dloneed byiMods Schs'dt
0
Active Materiats and Structures Laboratory
MIT
Dmt& "prt
Intended naruPfcturer' Aliflex
L"q, 1999
Ccersnt
and detalS
L Lines are Ia. eapper,
0
CD
ii
CIT
p1
CD
CIT
CD
MSl
wides
a. Lines separated by 0020'.
3 Solder pads are 10' square, and are an top surface
as seen In this drawbig.
of the flex clAt
4. The Lnes flare out at one end of each layer. The Unes
are D13T wide, with C0LS0' between them
* The autine of each layer Indicates where the kapton should
lwe cut Toierane.
shoad be as tlWt
ae possibl*.
* The "apton iould be 0.101' tNek, and should cover bc4t, side.
of the flexcircalt, except for the top
pads.
side
of
the
square solder
d
7-77
1
1
1
1.620"/
1-
-0 0.250 "
I
1.0 0"
0,050
R. T y p.
1.000"1
3/8
endrimil
occept a be
rLL
--
- 1,820"
Figure B-4: Drawing of outboard tip fitting. Material is steel.
157
1, 9 1"
(14
5.773"/
Linewidth: 0.007"
Line spacing: 0.045"
2.610
0,500/"( T yp,)
0.500 "(T y p.)
Figure B-5: Section 4.0 electrode pattern. The design is based on the Rodgers packs [Rodgers,
1998].
158
Appendix C
AMR Design and Testing
Supplement
C.1
Blade Actuator Pack Data
The actuator packs that were used in the process blade are shown in Table C.1.
C.2
Bending Stiffness
Figure C-1 shows the average bending stiffness of the blade, as detailed in Section 5.3.3.
C.3
Blade Dynamics
Transfer functions were gathered during testing of the AMR blade.
Two transfer functions
were found: from drive voltage in twist mode to twist rate in the active area, and from drive
voltage in bending mode to curvature of the active area. The natural frequencies of the process
blade are:
" 2nd bending: 37.5Hz
" 1st torsion: 91.0Hz
" 3rd bending: 102.0Hz
159
Pack Number
Location in Blade
Actuation (2KVpp/3KVpp)
007
010
011
018
025
030
008
009
014
017
022
029
1
2
3
4
5
6
7
8
9
10
11
12
800/1500
900/1600
750/1300
715/1280
700/1225
700/1250
750/1400
750/1350
735/1320
710/1270
750/1350
675/1225
744/1339
Average
Table C.1: Actuation data for process blade packs.
200
-
180160-140
120100 < 80
60 40 20 -
0
L
0
5
15
20
10
Distance from root clamp (in)
25
30
Figure C-1: Average bending stiffness between root clamp and given measurement location.
160
*
4th bending: 197.5Hz
The predicted natural frequencies, using Boeing dynamic models1 , for a clamped root, were:
* 2nd bending: 41.1Hz
" 1st torsion: 88.5Hz
* 3rd bending: 105.9Hz
" 4th bending: 209.0Hz
A scanning laser vibrometer 2 is used to find the average spectrum of the displacement of
several points along the span of the blade. See Figures C-4 and C-5.
Graphical representations of the mode shapes of the AMR blade are shown in Figures C-6
to C-13.
These mode shapes were obtained using the scanning laser vibrometer.
The drive
levels used during vibrometer testing are very low due to velocity measurement limitations of
the vibrometer.
The exact drive levels are not known. Therefore, the information provided in
Figure C-6 to C-13 should be used for visualization purposes only.
C.4
Blade Crossectional Pictures
The AMR blade was crossectioned after testing was completed.
A diamond abrasive rotary
say was used. A feed rate of 3.6" per minute was used to ensure a good surface finish. Figures
C-14 and C-15 show the complete set of blade crossections, indicating the blade station at which
each was cut.
C.5
AMR Flexible Circuit
The AMR blade flexible circuit is shown in Figures C-16 and C-17.
Figure C-16 shows the
drawing that was sent to the vendor, AllFlex, Inc 3 . Figure C-17 shows the blade station location
of each solder pad on the flexible circuit.
'Bobby Mathew, Boeing Helicopters, Philadelphia, PA.
Laser Vibroniter, Model PSV-300-F, Polytec PI, Inc. Tustin, CA
3AllFlex, Inc., Norfield, MN.
2Scanning
161
10-2
I
I
40
60
80
40
60
80
I
I
100
Freq (Hz)
120
140
160
100
Freq (Hz)
120
140
160
10-3
10-41
20
U
-0.5-
CL
-1. 5-
-220
Figure C-2: Transfer function from drive voltage to blade twist rate in the active area.
104
10
EU
10
-6
10 -71
40
60
100
80
120
140
14 0
Freq (Hz)
0
I
I
I
-0.5
-1
LO-1.5
0)
-2
-2.5
-3
-3.5
20
40
60
100
80
120
140
160
Freq (Hz)
Figure C-3: Transfer function from drive voltage (in bending mode) to curvature in the active
area.
162
"I., %-,-
I-I % ..
..........
Figure C-4: Average spectrum of the AMR blade displacement in twist mode.
Figure C-5: Average spectrum of the AMR blade displacement in bending mode.
163
x
Sig: Dis j
Figure C-6: Second bending mode of the AMR blade. The first mode is not shown since the
frequency of that mode is below the lower operating frequency of the amplifier.
Dwmn
FT
Figure C-7: Second bending mode of the AMR blade.
164
..
........
Figure C-8: Third bending mode of the AMR blade.
Figure C-9: Third bending mode of the AMR blade.
165
Z
X
Figure C-10: Fourth bending mode of the AMR blade.
VY
Figure C-11: Fourth bending mode of the AMR blade.
166
Figure C-12: First torsion mode of the AMR blade.
Figure C-13: First torsion mode of the AMR blade.
167
BS11.7
BS13.8
BS15.1
BS15.8
BS17.8
BS19.6
BS21.0
BS23.3
BS23.9
BS26.0
BS27.8
BS30.1
Figure C-14: Crossections of the AMR blade.
168
BS31.9
BS34.1
BS34.9
BS37.8
BS39.8
BS48.0
BS49.9
BS51.9
BS54.3
BS55.0
Figure C-15: Crossections of the AMR blade.
169
Layer 3
mo0
Layer
i
-
m
-
1
I
Emmomm"
.
-
--J
I
2815'
7.63f'
HIL
C)j
Layer 2
Eiiiiiiiiiing-
In
Layer 4
~p
-
m
1:
m
'-1
'Ear'
Det ait
150'
0
0.25'
R0.20'
DrIlL thru for 8-32 C170')-
D.0M0' thick G-10 repinforment; pod.
Pads should be 0.4' on each side,
centered on the through hole.
Mounted on top surface of circult,
as seen In drawing
Notes1) Solder pads are 0.00' on a side, and are exposed on top surface as seen In drawing,
Incorporate anchoring spurs on corners of soldering pads.
2) Lines are 0.010' wide, and are spaced 0.020' apart. loz. copper conductor.
3) 1.001' kapton covers both sides of flexcircult, except for top surface of square soldering pads.
4) ALL dimensions should fall out of drawing. For reference, the total
length Is 65472' (for longest cdrcults)
5) There is no exposed copper In 'flareouts' at the left hand side of drawing.
Only exposed copper Is the 0.l' square solder pads.
6) Outline of flexcircult indicates the trim Une of the flexclrcult.,
7) Please round sharp cornern In conductor Unes as In previous design
of MIT 4 Layer FLexclrcult, fron April 1999.
I
I
IF
UF
S
m
Aj
I
Figure C-17: Location of the solder pads in the AMR flexible circuit.
171
172
Appendix D
Flexible Circuit Connectors
12 pin connectors are crimped onto the flexible circuits to allow for interfacing to the power
distribution board. These connectors are female Berg Clincher connectors
The flexible circuit flareouts are cleaned with acetone to remove contaminants.
The flare-
outs are trimmed to match up with the spacing fo the female headers on the power distribution
board. These headers are 0.3" apart.
See Figure D-1.
The trim distances for the 4 layers are
as follows:
" Layer 1: 0"
" Layer 2: 0.3"
* Layer 3: 0.6"
" Layer 4: 0.9"
The connectors are then crimped onto the flareout, taking care to alight the crimps with
the conductors in the flexible circuit.
See Figure D-2.
Continuity is checked from each solder
pad in the circuit to the corresponding crimp connection.
connections.
See Figure D-3.
Solder is applied to all the crimp
This causes solder to flow acound the crimp and onto the
copper of the flexible circuit line. This ensures a better electrical contact. Again, continuity is
checked. Straight pins (such as from as straight header) are inserted into the receptacles on the
'Berg Clincher connector. Newark Electronics P/N: 89F4613. Type: 65801-012.
173
female connectors.
This converts the connectors from female to male, as shown in Figure D-4.
Using male clincher connectors would prevent removal and replacement of the pins following
the potting procedure.
Using female connectors and inserting pins into the receptacle allows
individual lines to be disconnected and reconnected easily by simply removing and replacing
pins.
Continuity is checked from each pin to the corresponding solder pad.
Vacuum tape or other thick, sticky material is placed around the pins of the connector to
prevent epoxy from coating the pins. A mold around the connector is made with GNPT tape.
This mold will allow epoxy to be poured around the connector to perform the potting.
The
mold is shown in Figure D-5.
Shell 828 epoxy resin with 3223 curing agent is used 2 . The mix ratio is 100:11.
air release 3 is added to the epoxy to ensure that trapped air bubbles are removed.
A drop of
The epoxy
is poured slowly into the GNPT mold, which is oriented vertically as shown in Figure D-6. The
mold should be inspected to ensure that no large trapped bubbles are visible.
The epoxy will
cure at room temperature in approximately 8 hours.
Following the cure, the GNPT and vacuum tape is removed. The flexible circuits are tested
by plugging them into a power distribution board and testing them up to 5000Vpp. The design
voltage is 3000Vpp. If no failures occur, the flexible circuit is considered functional. If failures
do occur in the connector, the new connector must be installed.
2
Shell 828 resin, 3223 curing agent, Miller-Stephenson, Danbury, CT., 203-743-4447.
3Air release agent A530,
BYK-Chemie, Wallingford, CT.
174
..
. ..
.......
Figure D-1: The flexible circuit flareouts are trimmed to match the stagger on the power
distribution board. Layer 1 is the longest, and Layer 4 is the shortest.
175
Figure D-2: The Berg Clincher connectors are crimped onto the trimmed flareout.
flap of the connector is removed.
Figure D-3: Solder is applied to each crimp of the connector.
nection to the copper lines in the flareout.
176
The top
This improves the crimp con-
receptacles on the Clincher connectors.
Figure D-4: Header pins are inserted into
Figure D-5: Vacuum tape is applied to cover the pins and create the lower portion of the potting
mold. GNPT tape is wrapped around the vacuum tape to finish the mold.
177
Figure D-6: The molds are mounted vertically so that the potting material (epoxy) can be
poured into the mold.
178
Appendix E
Piezoelectric Ceramic Data
This appendix contains data on the electrical properties of PZT-5A and PZT-4.
PZT-4 is a
harder ceramic, with lower actuation levels than PZT-5A. However, PZT-4 electrical properties
do not increase as rapidly with voltage as those of PZT-5A.
Table E.1 shows some relevant properties of PZT-5A, PZT-4, and PZT-8.
This data was
obtained from [Berlincourt, 2000].
Description
PZT-5A
PZT-4
PZT-8
d3 3
sE
Piezoelectric constant (1Ou /)
Mechanical compliance (1O m 2 )
374
18.8
289
15.5
225
13.5
p
Density (103 k)
7.75
7.5
7.6
Table E.1: Relevant piezoelectric properties of PZT-5A, PZT-4, and PZT-8
179
40
PZT-5
30
00
PPZT-4
20
10
CO
PZT-8
0
0.08
PZT-5A
0.06
PZT-4
0.04
0.02
PZT-8
0
1
2
4
3
5
AC RMS Field, kV/cm
6
7
Figure E-1: E3 and tan 8 information on PZT-5A, PZT-4, and PZT-8. The dielectric constant
and the dielectric loss factor of PZT-5A increase much more rapidly with field than PZT-4.
This data was obtained from [Berlincourt, 2000b].
180
Appendix F
Core Void Elimination
The cause of numerous large voids in the spar of the blade was studied. It was shown that
one of the chemicals used in applying strain gages to the spar core outgassed during the cure,
forming the large voids.
F.1
Motivation
Destructive damage analysis was performed on the first generation integral blade upon completion of hover testing. The blade was cut into sections and the crossections were inspected
for damage or defects. Figure F-1 shows a map of the damage observed in the first generation
blade. Large voids were observed in the core of the spar. A typical void is shown in Figure F-2.
A complete set of crossectional pictures have been published [Rodgers, 19981.
The void locations correlated well with the locations of the embedded strain gages. Of the
11 embedded strain gage bridges, 10 had voids in the same location or in the vicinity of the
gages. No significant core voids were found in locations not in the vicinity of strain gages.
It was hypothesized that the core voids were a result of either heat, chemical interactions,
or trapped air or moisture. [Rodgers, 1998].
Large core voids are obviously undesirable. Voids can cause strain gage wires to separate
from the foam core. This would put them at risk for breaking off from the strain gages due
to lack of proper support. Large voids under the laminate also remove the support for the
laminate, and can lead to delamination of the laminate during actuation.
181
- - - ---
2 )9' 1 X1 4 X
-
section cut
o
strain gage
>
delamination (inner)
<
delamination (outer)
*
pack electrical failure
>
ebctrode bman
UPPER
L--------
+
core void
"
BI 9.093
17
21
25
29
33
37
--J
--------------41
45
49
53
8
57 60. 619
ncore
4
1
17
i1
5
2
3
7
41
45
49 13
57 6C. 619
LOWER
Figure F-1: Map of damage found in the first generation blade upon completion of hover testing.
F.2
Hypotheses of void formation
Two main hypotheses for the void formation were pursued in the study. The first hypothesis
focuses on the adhesive used to attach the strain gages to the spar core. The second focused
on the role of moisture trapped in the foam in the formation of the voids.
Before the layup of the blade, the spar core is instrumented with several strain gage bridges.
The strain gages are mounted to the spar core using M-Bond 200 adhesive'. The M-Bond 200
adhesive consists of two parts. The first part is the adhesive itself, and the second is a catalyst.
The adhesive is applied by first coating the part to be bonded with a thin coat of catalyst.
1
M-Bond 200, Micromeasurements Group, Raleigh, NC.
Figure F-2: Crossection of first generation blade at BS19.
uppper portion of spar core.
182
Spar core void clearly visible in
The part that is to be bonded to is then coated with the adhesive and the strain gage is then
lowered into place.
It was hypothesized that the M-Bond 200 adhesive was to blame in the formation of the
voids. This hypothesis was based mainly on the good correlation between strain gage location
and the location of the voids.
Also, M-Bond 200 is a cyanoacrylate adhesive, a family of
adhesives that are notoriously difficult to cure correctly.
Moisture was also hypothesized as a possible cause of the voids. The Rohacell 2 foam used in
the spar core is notorious for moisture absorption. Since the foam was not heat treated before
the layup, the moisture in the foam may have evaporated, forming the voids seen in the core.
The correlation with the strain gage locations may have been due to the strain gages acting as
initiation sites for the moisture-induced voids.
F.3
Test procedure
To test the hypotheses described above, a spar section was manufactured. This spar section
was built to simulate the blade cure. The manufacturing procedure was the same as was used
in the first generation integral blade. The foam core, or mandrel, of this spar consisted of six
5" long blocks of foam. These foam blocks were shaped on a belt sander using steel templates
to achieve the correct shape.
The blocks were numbered 1 through 6. Blocks 1, 2, and 3 were baked for 90 minutes at
250'F prior to attachment of strain gages. This was done to remove any moisture content in
the foam. The foam, when stored, was kept in a vacuum bag to prevent moisture absorption.
On the spar mandrel, 8 strain gages were mounted following the Boeing strain gaging
procedure 3 . The strain gages were mounted to a piece of E-glass with AE-15 epoxy4 . Strain
gaging wire was then soldered to the strain gages. Four of these strain gage assemblies were
then mounted to the top surface of the mandrel using M-Bond 200 adhesive as shown in Figure
F-3. Two strain gages were mounted on the baked foam and two on the non-baked foam. A
significant amount of adhesive was necessary to make the bond because some of the adhesive was
2
Rohacell, Richmond Aircraft, Norwalk, CA.
Richard Bussom, Boeing Helicopters, Philadelphia, PA.
4AE-15 epoxy, Micromeasurements Group, Raleigh, NC.
3
183
.
...
................
.......
...
.......
......
. ............
5
6
Figure F-3: Foam blocks assembled into spar core. Strain gages mounted on blocks 2 through 5,
on upper and lower surfaces. Upper gages mounted with M-Bond 200 and lower gages mounted
with 5-minute epoxy.
absorbed by the foam. Four strain gages were mounted to the bottom surface of the mandrel
5
using generous amounts of 5-minute epoxy . Two gages were mounted on the baked foam, and
two were mounted to the non-baked foam.
Following the mounting of the strain gages, the laminate was laid up around the spar
mandrel. The laminate that was used in shown in Figure F-4. The thickness of the laminate is
similar to the thickness of the blade laminate. Two plies of Kapton were incorporated to ensure
5
Devcon 5-Minute Epoxy, Danvers, MA.
E-glass @ 45 deg
E-glass @ 45 deg
S-glass @ 0 deg
....
E-glass @ 45 deg
E-glass @ 45 deg
Kapton
E-glass @ 45 deg
Kapton
E-glass @ 45 deg
IM7 Graphite @ 0 deg
Figure F-4: Layup used on the test spar. Kapton is included in the layup to render it impermeable to gas.
184
Figure F-5: Spar cut into 1" sections using a diamond abrasive rotary saw.
impermeability to gas and to simulate the AFC pack surfaces.
Following the layup, the test piece was cured in the blade mold at 250'F for 90 minutes,
the standard cure cycle for the blades. After the blade cure, the spar was removed from the
mold and cut into 1" sections using a diamond abrasive
circular saw. A schematic of the
crossectioning cuts is shown in Figure F-5.
F.4
Results
The crossections of the spar section were analyzed and inspected for voids. Several voids were
found throughout the spar. Figure F-6 shows the location and extent of the observed voids.
Large voids were found on top and around the strain gages on the upper surface of blocks
2, 3, and 4. No voids were observed around the gage mounted on block 5.
Slight depressions
in the foam were seen where the bottom surface gages were attached. Five small voids were
seen in blocks 4, 5, and 6. No delaminations were seen in the laminate around the voids.
Figure F-7 is a typical crossection of the spar section. The entire set of crossections have
been included in Section F.7.
185
2
6
3
5
4
4
5
3
6
2
Figure F-6: Top and bottom surfaces of spar mandrel. Showing locations of strain gages and
voids. Voids shown as dark ovals. Small depressions in the foam were also observed where the
bottom surface gages were tacked onto the foam (not shown in this picture due to small size).
F.5
Analysis
Figure F-6 shows the location of the voids that were observed in the spar. It can be seen that
large voids were found at the location of three of the four strain gages. These voids were similar
in shape and size to the voids observed in the phase I blade. Figure F-8 shows a detailed view
of a void. The groove in which the strain gage wire was located has been pushed away from
the laminate. This indicates that the foam was pushed away from the laminate, rather than
being dissolved or evaporated. If the foam had been chemically dissolved, the strain gage wire
groove would not have survived as a sharp indentation. Figure F-9 shows crossections from the
test next to crossections from the Phase I blade. A clear resemblance can be seen.
During the manufacturing procedure of both the Phase I blade and the test piece, excess
M-Bond 200 adhesive was used when attaching the strain gages to the core foam.
This was
necessary because the adhesive seeped into the foam, making it difficult to achieve a good bond
between the foam and the strain gage. Since only a very small amount of catalyst was applied
186
Figure F-7: Crossectional picture of sections cut from block 4. A large void can be observed in
the top surface of the mandrel. A smaller void can be seen near the web in one of the sections.
This void does not correlate with strain gage location.
to the strain gage before bonding, the adhesive that seeped into the foam may not have come in
contact with the adhesive, and may not have cured properly. During the elevated temperature
and pressure of the blade cure, the adhesive may have outgassed or evaporated, pushing the
foam away from the gage, thereby creating voids in the relatively soft spar core.
The lack of a void around the strain gages mounted with M-Bond 200 on block 5 was most
likely due to a proper curing of the adhesive.
Enough catalyst may have been used on this
gage to completely cure the M-Bond 200 adhesive. Cyanoacrylate adhesive is typically difficult
to cure properly because very exact ratios of catalyst to adhesive are necessary.
The small voids observed around the strain gages mounted with 5-minute epoxy appear to
have been caused by mechanical deformation of the foam due to the thick layer of epoxy under
each strain gage. When the epoxy cured, it formed a hard layer under the strain gage. When
the blade molds were closed, this hard layer of epoxy was pressed into the foam, mechanically
deforming it. At the elevated temperatures, the 5 minute epoxy softened and may have flowed,
leaving a void where the hard epoxy layer had pressed into the foam.
Small voids were observed in the foam sections that had not been dried prior to the blade
cure. These voids can be seen in Figure F-10. The location of the small voids does not correlate
with strain gages or any other embedded object or chemical. These voids may therefore be
187
Figure F-8: Detailed view of an adhesive-caused void. The void is similar to the voids found in
the Phase I blade.
attributed to outgassing of the foam itself.
The lack of moisture in the dried foam further
reinforces this point. These voids are not present in the foam that had been dried prior to the
blade cure.
F.6
Recommendations
The blade manufacturing procedure should be changed to avoid the formation of voids in the
blade core. First, all moisture should be removed from the core foam prior to the blade cure.
This can be done by baking the foam for approximately one hour at the cure temperature of
the blade, 250'F. The foam should then be kept in sealed plastic bags with dessecant packs
until the layup of the blade is performed. This will prevent moisture from being absorbed into
the foam between baking and the blade cure.
M-Bond 200 should be removed from the manufacturing process.
Since the M-Bond 200
adhesive was only used to make a non-critical bond between the strain gage and the foam (the
gage is designed to be cured to the inner surface of the spar laminate), epoxy may be substituted
for the M-Bond. In the void formation testing, a large amount of 5 minute epoxy was used, and
no significant voids were observed. However, to reduce the deformation of the foam underneath
the strain gage, only the corners of the gage should be tacked down with epoxy.
188
Figure F-9: Picture showing Phase I blade voids (left) and testpiece voids (right).
resemblance can be seen in void shape, size, and character.
F.7
A clear
Voids Formation Testing Details
The following pictures are the crossectional pictures of the spar section. The crossections are
numbered starting from the outboard end of the spar section. The crossections that correspond
to each foam block are grouped and are shown in the same picture. Both the inboard and
outboard sides of each section piece was photographed and is presented.
189
Figure F-10: Small voids were found in the non-dried sections of foam. These voids can be
attributed to moisture trapped in the foam. During the cure, the moisture is released and
causes core voids.
Figure F-11: Foam block 1: Crossections 1-4, outboard view.
190
Figure F-12: Foam block 1: Crossections 1-4, inboard view.
Figure F-13: Foam block 2: Crossections 5-9, outboard view.
Figure F-14: Foam block 2: Crossections 5-9, inboard view.
191
Figure F-15: Foam block 3: Crossections 10-14, outboard view.
Figure F-16: Foam block 3: Crossections 10-14, inboard view.
Figure F-17: Foam block 4: Crossections 15-20, outboard view.
192
Figure F-18: Foam block 4: Crossections 15-20, inboard view.
Figure F-19: Foam block 5: Crossections 21-25, outboard view.
Figure F-20: Foam block 5: Crossections 21-25, inboard view.
193
......
- .-,.....
.....
...
. ........
Figure F-21: Foam block 6: Crossections 26-29, outboard view.
Figure F-22: Foam block 6: Crossectiong 26-29, inboard view.
194
195