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AN ABSTRACT OF A THESIS
J-INTEGRAL FINITE ELEMENT ANALYSISOF SEMI-ELLIPTICAL SURFACE
CRACKS IN FLAT PLATESWITH TENSILE
LOADING
Eric N. Quillen
Master of Science in Mechanical Engineering
Linear elastic fracture mechanics (LEFM) is used when response to the loadis elastic, and the fracture is brittle. For LEFM, the K-factor is the most commonlyused fracture criterion. However, high temperatures and limited high stress cycles be-fore component replacement are factors that can cause significant plastic deformationand a ductile failure. In these cases, an elastic-plastic fracture mechanics (EPFM)approach is required. The J-integral is commonly used as an EPFM fracture param-eter.
The primary goal of this research was to develop three-dimensional finite el-ement analysis (FEA) J-integral data for surface crack specimen geometries andcompare to existing solutions. The finite element models were analyzed as elas-tic, and fully plastic using ABAQUS. The J-integral data were used to find the loadindependent variable, h1 for comparison purposes.
There were two other goals in this research. The second goal was to examinethe effect of various finite element modelling parameters including mesh density, ele-ment type, symmetry, and specimen size effects, on the resulting J-integral. The thirdgoal was to perform elastic-plastic finite element analyses that utilize a stress vs. plas-tic strain table based on a power law hardening material behavior. The elastic-plasticand fully plastic results were compared.
For the most part, the current data compared well with the data published byother researchers. The elastic results compared more favorably than the fully plasticand elastic-plastic data. For both the elastic and plastic analyses, the finite elementmodels (FEMs) produced sudden increases in the K-factor and J-integral at the freesurface and/or depth. The plastic FEMs also exhibited an anomaly in the J-integralat the third and fourth angles from the surface. The anomaly could be taken as ajump at the third angle or a dip at the fourth angle, depending on how the data weretrended. The third angle varied with the model geometry (2.71◦ to 11.24◦).
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J-INTEGRAL FINITE ELEMENT ANALYSIS
OF SEMI-ELLIPTICAL SURFACE
CRACKS IN FLAT PLATES
WITH TENSILE
LOADING
A Thesis
Presented to
the Faculty of the Graduate School
Tennessee Technological University
by
Eric N. Quillen
In Partial Fulfillment
of the Requirements for the Degree
MASTER OF SCIENCE
Mechanical Engineering
May 2005
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STATEMENT OF PERMISSION TO USE
In presenting this thesis in partial fulfillment of the requirements for a Master
of Science degree at Tennessee Technological University, I agree that the University
Library shall make it available to borrowers under rules of the Library. Brief quota-
tions from this thesis are allowable without special permission, provided that accurate
acknowledgment of the source is made.
Permission for extensive quotation from or reproduction of this thesis may be
granted by my major professor when the proposed use of the material is for scholarly
purposes. Any copying or use of the material in this thesis for financial gain shall not
be allowed without my written permission.
Signature
Date
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DEDICATION
This thesis is dedicated to my wife Julie, whose encouragement has been critical
in the completion of my graduate degree and the composition of this thesis.
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ACKNOWLEDGMENTS
I would like to thank the following people for their help with this work: Dr.
Chris Wilson, Dr. Phillip Allen, Mike Renfro, Krishna Natarajan, and Richard
Gregory. I would also like to thank my employer, Fleetguard, Inc., and cowork-
ers. Without their cooperation, it would not have been possible for me to perform
this research.
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TABLE OF CONTENTS
Page
LIST OF TABLES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ix
LIST OF FIGURES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xiii
LIST OF SYMBOLS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . xx
Chapter
1. INTRODUCTION . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1
1.1 Fracture Mechanics . . . . . . . . . . . . . . . . . . . . . . . 1
1.2 Overview of Research . . . . . . . . . . . . . . . . . . . . . . 2
2. TECHNICAL BACKGROUND . . . . . . . . . . . . . . . . . . . . . . 4
2.1 J-Integral . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4
2.2 EPRI Estimation Scheme . . . . . . . . . . . . . . . . . . . . 8
2.3 Reference Stress Method . . . . . . . . . . . . . . . . . . . . 17
3. RESEARCH PROCEDURE . . . . . . . . . . . . . . . . . . . . . . . . 19
3.1 Finite Element Modeling . . . . . . . . . . . . . . . . . . . . 19
3.1.1 Mesh Generation . . . . . . . . . . . . . . . . . . . . . . 21
3.1.1.1 mesh3d scp . . . . . . . . . . . . . . . . . . . . . . . 21
3.1.1.2 FEA-Crack . . . . . . . . . . . . . . . . . . . . . . . 22
3.2 Analysis Procedure . . . . . . . . . . . . . . . . . . . . . . . 24
3.3 J-Integral Convergence . . . . . . . . . . . . . . . . . . . . . 26
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Chapter Page
3.3.1 Load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26
3.3.2 Fully Plastic Zone . . . . . . . . . . . . . . . . . . . . . 27
3.4 Comparison to Other Work . . . . . . . . . . . . . . . . . . . 31
3.4.1 Kirk and Dodds . . . . . . . . . . . . . . . . . . . . . . . 31
3.4.2 McClung et al. [15] . . . . . . . . . . . . . . . . . . . . . 35
3.4.3 Lei [17] . . . . . . . . . . . . . . . . . . . . . . . . . . . 37
3.4.4 Nasgro Computer Program . . . . . . . . . . . . . . . . 38
3.5 Mesh Refinement . . . . . . . . . . . . . . . . . . . . . . . . 40
3.6 Finite Size Effects . . . . . . . . . . . . . . . . . . . . . . . . 40
3.7 Material Properties . . . . . . . . . . . . . . . . . . . . . . . 41
3.7.1 Deformation Plasticity . . . . . . . . . . . . . . . . . . . 41
3.7.2 Incremental Plasticity . . . . . . . . . . . . . . . . . . . 43
4. RESULTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49
4.1 Fully Plastic Zone . . . . . . . . . . . . . . . . . . . . . . . . 49
4.2 Kirk and Dodds Incremental Plasticity . . . . . . . . . . . . 49
4.3 McClung and Lei Comparisons . . . . . . . . . . . . . . . . . 50
4.3.1 Elastic Analysis . . . . . . . . . . . . . . . . . . . . . . . 51
4.3.2 Fully Plastic Analysis . . . . . . . . . . . . . . . . . . . 67
4.3.3 Incremental Elastic-Plastic Analysis . . . . . . . . . . . . 86
4.4 Mesh Refinement . . . . . . . . . . . . . . . . . . . . . . . . 91
4.5 Size Effects . . . . . . . . . . . . . . . . . . . . . . . . . . . 93
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Chapter Page
4.5.1 Height Effects . . . . . . . . . . . . . . . . . . . . . . . . 93
4.5.2 Width Effects . . . . . . . . . . . . . . . . . . . . . . . . 95
5. CONCLUSIONS AND RECOMMENDATIONS . . . . . . . . . . . . . 104
5.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . 104
5.2 Recommendations . . . . . . . . . . . . . . . . . . . . . . . . 106
REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108
APPENDICES
A: INSTRUCTIONS FOR MESH3D SCP MODIFICATIONS . . . . . . . . . 113
B: COARSE VERSUS REFINED MESHES FOR K-FACTORS . . . . . . . 115
C: COARSE VS. REFINED MESHES FOR FULLY PLASTIC MODELS . . 120
D: HEIGHT EFFECTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133
E: K-FACTOR RESULTS FOR COARSE MESHES . . . . . . . . . . . . . . 137
F: FULLY PLASTIC RESULTS FOR COARSE MESHES . . . . . . . . . . . 152
G: INCREMENTAL PLASTICITY TABLES . . . . . . . . . . . . . . . . . . 163
VITA . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 170
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LIST OF TABLES
Table Page
2.1 McClung et al. h1 values in tension, n = 15 [15] . . . . . . . . . . . . . 14
2.2 McClung et al. h1 values in tension, n = 10 [15] . . . . . . . . . . . . . 14
2.3 McClung et al. h1 values in tension, n = 5 [15] . . . . . . . . . . . . . . 15
2.4 Lei h1 values in tension, n = 5 [17] . . . . . . . . . . . . . . . . . . . . 16
2.5 Lei h1 values in tension, n = 10 [17] . . . . . . . . . . . . . . . . . . . . 16
3.1 Number of nodes and elements in the duplication of the Kirk and Dodds[23] geometries . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
3.2 Incremental plasticity values for the Kirk and Dodds models . . . . . . 33
3.3 McClung et al. fully plastic geometries . . . . . . . . . . . . . . . . . . 36
3.4 Geometries for Nasgro comparison and width effect investigation . . . . 39
3.5 Number of crack front nodes in the coarse and refined meshes . . . . . 40
3.6 Stress vs. plastic strain data at n = 15, used for ABAQUS models . . . 47
3.7 Stress vs. plastic strain data at n = 10, used for ABAQUS models . . . 47
3.8 Stress vs. plastic strain data at n = 5, used for ABAQUS models . . . 48
4.1 Comparison of FEM results to Kirk and Dodds values . . . . . . . . . . 50
4.2 Surface and depth phenomenon for K-factors . . . . . . . . . . . . . . 56
4.3 Maximum percent differences between Newman-Raju and FEM solutions(quarter symmetry) . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57
4.4 Maximum percent differences between McClung et al. [15] and FEMsolutions (quarter symmetry) . . . . . . . . . . . . . . . . . . . . . . 82
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Table Page
4.5 Maximum percent differences between McClung et al. [15] and Lei [17]solutions (quarter symmetry) . . . . . . . . . . . . . . . . . . . . . . 83
4.6 Model 1 (a/t=0.2 and a/c=0.2): h1 values at different heights . . . . . 94
4.7 Comparison of Nasgro and FEM results for n = 15 . . . . . . . . . . . 95
4.8 Comparison of Nasgro and FEM results for n = 10 . . . . . . . . . . . 96
4.9 Comparison of Nasgro and FEM results for n = 5 . . . . . . . . . . . . 96
D.1 Model 4 (a/t=0.5 and a/c=0.2): h1 values for at different heights (Part 1) 134
D.2 Model 4 (a/t=0.5 and a/c=0.2): h1 values for at different heights (Part 2) 135
D.3 Model 5 (a/t=0.5 and a/c=0.6): h1 values for at different heights . . . 135
D.4 Model 9 (a/t=0.8 and a/c=1.0): h1 values for at different heights . . . 136
E.5 Model 1 (a/t=0.2, a/c=0.2): K-Factor data from ABAQUS (Part 1) . 138
E.6 Model 1 (a/t=0.2, a/c=0.2): K-Factor data from ABAQUS (Part 2) . 139
E.7 Model 2 (a/t=0.2, a/c=0.6): K-Factor data from ABAQUS . . . . . . 140
E.8 Model 3 (a/t=0.2, a/c=1.0): K-Factor data from ABAQUS . . . . . . 141
E.9 Model 4 (a/t=0.5, a/c=0.2): K-Factor data from ABAQUS (Part 1) . 142
E.10 Model 4 (a/t=0.5, a/c=0.2): K-Factor data from ABAQUS (Part 2) . 143
E.11 Model 5 (a/t=0.5, a/c=0.6): K-Factor data from ABAQUS . . . . . . 144
E.12 Model 6 (a/t=0.5, a/c=1.0): K-Factor data from ABAQUS . . . . . . 145
E.13 Model 7 (a/t=0.8, a/c=0.2): K-Factor data from ABAQUS (Part 1) . 146
E.14 Model 7 (a/t=0.8, a/c=0.2): K-Factor data from ABAQUS (Part 2) . 147
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Table Page
E.15 Model 7 (a/t=0.8, a/c=0.2): K-Factor data from ABAQUS (Part 3) . 148
E.16 Model 8 (a/t=0.8, a/c=0.6): K-Factor data from ABAQUS (Part 1) . 149
E.17 Model 8 (a/t=0.8, a/c=0.6): K-Factor data from ABAQUS (Part 2) . 150
E.18 Model 9 (a/t=0.8, a/c=1.0): K-Factor data from ABAQUS . . . . . . 151
F.19 Model 1 (a/t=0.2, a/c=0.2): h1 data from ABAQUS . . . . . . . . . . 153
F.20 Model 2 (a/t=0.2, a/c=0.6): h1 data from ABAQUS . . . . . . . . . . 154
F.21 Model 3 (a/t=0.2, a/c=1.0): h1 data from ABAQUS . . . . . . . . . . 155
F.22 Model 4 (a/t=0.5, a/c=0.2): h1 data from ABAQUS (Part 1) . . . . . 156
F.23 Model 4 (a/t=0.5, a/c=0.2): h1 data from ABAQUS (Part 2) . . . . . 157
F.24 Model 5 (a/t=0.5, a/c=0.6): h1 data from ABAQUS . . . . . . . . . . 157
F.25 Model 6 (a/t=0.5, a/c=1.0): h1 data from ABAQUS . . . . . . . . . . 158
F.26 Model 7 (a/t=0.8, a/c=0.2): h1 data from ABAQUS (Part 1) . . . . . 159
F.27 Model 7 (a/t=0.8, a/c=0.2): h1 data from ABAQUS (Part 2) . . . . . 160
F.28 Model 8 (a/t=0.8, a/c=0.6): h1 data from ABAQUS . . . . . . . . . . 161
F.29 Model 9 (a/t=0.8, a/c=1.0): h1 data from ABAQUS . . . . . . . . . . 162
G.30 Stress vs. strain data at n = 15, based on Equation 3.13 . . . . . . . . 164
G.31 Stress vs. strain data at n = 10, based on Equation 3.13 . . . . . . . . 165
G.32 Stress vs. strain data at n = 5, based on Equation 3.13 . . . . . . . . . 166
G.33 Stress vs. plastic strain data at n = 15, used for ABAQUS models . . . 167
G.34 Stress vs. plastic strain data at n = 10, used for ABAQUS models . . . 168
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G.35 Stress vs. plastic strain data at n = 5, used for ABAQUS models . . . 169
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LIST OF FIGURES
Figure Page
2.1 Contour around a crack tip [4] . . . . . . . . . . . . . . . . . . . . . . . 5
2.2 EPRI J-Integral estimation scheme [4] . . . . . . . . . . . . . . . . . . 9
2.3 Sample of finite element mesh used by McClung et al. [15] . . . . . . . 12
2.4 Close up of the finite element mesh around the crack front used byMcClung et al. [15] . . . . . . . . . . . . . . . . . . . . . . . . . . . 12
3.1 Degeneration of elements around crack tip [4] . . . . . . . . . . . . . . 20
3.2 Plastic singularity element [4] . . . . . . . . . . . . . . . . . . . . . . . 20
3.3 Zones created in the mesh by mesh3d scp [20] . . . . . . . . . . . . . . 22
3.4 Mesh created using FEA-Crack . . . . . . . . . . . . . . . . . . . . . . 23
3.5 Close up of mesh from Figure 3.4 created using FEA-Crack . . . . . . . 23
3.6 Contours (semi-circular rings) around the crack tip . . . . . . . . . . . 24
3.7 Coordinate scheme for mapping crack face angles . . . . . . . . . . . . 26
3.8 Fully plastic element set consisting of the elements around the crack tip 28
3.9 Fully plastic element set consisting of part of layer 1 . . . . . . . . . . . 29
3.10 Fully plastic element set consisting of layer 1 . . . . . . . . . . . . . . . 29
3.11 Fully plastic element set consisting of partial layers 1 and 2 . . . . . . . 30
3.12 Geometries used by Kirk and Dodds for estimating the J-Integral [23] . 32
3.13 Stress vs. strain curve for Kirk and Dodds elastic-plastic models [23] . . 34
3.14 Refined mesh along the crack front . . . . . . . . . . . . . . . . . . . . 41
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Figure Page
3.15 Effect of n on the stress vs. strain curve using a Ramberg-Osgood model 42
3.16 Intersection of Ramberg-Osgood curves at σo . . . . . . . . . . . . . . . 44
3.17 Elastic, modified elastic, and Ramberg-Osgood stress vs. strain curves forn = 10 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45
4.1 Model 1 (a/t=0.2, a/c=0.2): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51
4.2 Model 2 (a/t=0.2, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52
4.3 Model 3 (a/t=0.2, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52
4.4 Model 4 (a/t=0.5, a/c=0.2): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
4.5 Model 5 (a/t=0.5, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53
4.6 Model 6 (a/t=0.5, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54
4.7 Model 7 (a/t=0.8, a/c=0.2): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54
4.8 Model 8 (a/t=0.8, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
4.9 Model 9 (a/t=0.8, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
4.10 Elastic singularity element [4] . . . . . . . . . . . . . . . . . . . . . . . 58
4.11 Model 1 (a/t=0.2, a/c=0.2): Normalized K-factor vs. angle along crackfront for untied and tied nodes . . . . . . . . . . . . . . . . . . . . . 58
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Figure Page
4.12 Model 8 (a/t=0.8, a/c=0.6): Normalized K-factor vs. angle along crackfront for untied and tied nodes . . . . . . . . . . . . . . . . . . . . . 59
4.13 Model 1 (a/t = 0.2, a/c = 0.2): Reduced vs. full integration elements . 60
4.14 Model 8 (a/t = 0.8, a/c = 0.6): Reduced vs. full integration elements . 61
4.15 K-factor results from FEA-Crack Validation Manual [26] . . . . . . . . 63
4.16 FEM mesh for a flat plate with no symmetry exploited [26] . . . . . . . 63
4.17 FEM mesh for a flat plate with half symmetry . . . . . . . . . . . . . . 64
4.18 Model 6 (a/t = 0.5, a/c = 1.0): K-Factor results for half symmetry model 64
4.19 Model 8 (a/t = 0.8, a/c = 0.6): K-Factor results for half symmetry model 65
4.20 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle along the crack front . . . . . 67
4.21 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle along the crack front . . . . . 68
4.22 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle along the crack front . . . . . 68
4.23 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle along the crack front . . . . . 69
4.24 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle along the crack front . . . . . 69
4.25 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle along the crack front . . . . . 70
4.26 Model 3 (a/t=0.2, a/c=1.0): h1 vs. angle along the crack front . . . . . 70
4.27 Model 3 (a/t=0.2, a/c=1.0): h1 vs. angle along the crack front . . . . . 71
4.28 Model 3 (a/t=0.2, a/c=1.0): h1 vs. angle along the crack front . . . . . 71
4.29 Model 4 (a/t=0.5, a/c=0.2): h1 vs. angle along the crack front . . . . . 72
4.30 Model 4 (a/t=0.5, a/c=0.2): h1 vs. angle along the crack front . . . . . 72
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4.31 Model 4 (a/t=0.5, a/c=0.2): h1 vs. angle along the crack front . . . . . 73
4.32 Model 5 (a/t=0.5, a/c=0.6): h1 vs. angle along the crack front . . . . . 73
4.33 Model 5 (a/t=0.5, a/c=0.6): h1 vs. angle along the crack front . . . . . 74
4.34 Model 5 (a/t=0.5, a/c=0.6): h1 vs. angle along the crack front . . . . . 74
4.35 Model 6 (a/t=0.5, a/c=1.0): h1 vs. angle along the crack front . . . . . 75
4.36 Model 6 (a/t=0.5, a/c=1.0): h1 vs. angle along the crack front . . . . . 75
4.37 Model 6 (a/t=0.5, a/c=1.0): h1 vs. angle along the crack front . . . . . 76
4.38 Model 7 (a/t=0.8, a/c=0.2): h1 vs. angle along the crack front . . . . . 76
4.39 Model 7 (a/t=0.8, a/c=0.2): h1 vs. angle along the crack front . . . . . 77
4.40 Model 7 (a/t=0.8, a/c=0.2): h1 vs. angle along the crack front . . . . . 77
4.41 Model 8 (a/t=0.8, a/c=0.6): h1 vs. angle along the crack front . . . . . 78
4.42 Model 8 (a/t=0.8, a/c=0.6): h1 vs. angle along the crack front . . . . . 78
4.43 Model 8 (a/t=0.8, a/c=0.6): h1 vs. angle along the crack front . . . . . 79
4.44 Model 9 (a/t=0.8, a/c=1.0): h1 vs. angle along the crack front . . . . . 79
4.45 Model 9 (a/t=0.8, a/c=1.0): h1 vs. angle along the crack front . . . . . 80
4.46 Model 9 (a/t=0.8, a/c=1.0): h1 vs. angle along the crack front . . . . . 80
4.47 Model 6 (a/t=0.5, a/c=1.0): h1 results for half symmetry model at n = 15 84
4.48 Model 8 (a/t=0.8, a/c=0.6): h1 results for half symmetry model at n = 15 85
4.49 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle for fully plastic andelastic-plastic models at n=5 . . . . . . . . . . . . . . . . . . . . . . 87
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4.50 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle for fully plastic andelastic-plastic models at n=10 . . . . . . . . . . . . . . . . . . . . . . 87
4.51 Model 1 (a/t=0.2, a/c=0.2): h1 vs. angle for fully plastic andelastic-plastic models at n=15 . . . . . . . . . . . . . . . . . . . . . . 88
4.52 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle for fully plastic andelastic-plastic models at n=5 . . . . . . . . . . . . . . . . . . . . . . 88
4.53 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle for fully plastic andelastic-plastic models at n=10 . . . . . . . . . . . . . . . . . . . . . . 89
4.54 Model 2 (a/t=0.2, a/c=0.6): h1 vs. angle for fully plastic andelastic-plastic models at n=15 . . . . . . . . . . . . . . . . . . . . . . 89
4.55 Elastic, Ramberg-Osgood, modified elastic, and modifiedRamberg-Osgood stress vs. strain curves for n = 10 . . . . . . . . . . 90
4.56 Model 1 (a/t = 0.2, a/c = 0.2): Normalized K-factor vs. angle alongcrack front . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91
4.57 Model 1 (a/t = 0.2, a/c = 0.2): h1 vs. angle along the crack front . . . 92
4.58 Model 1 (a/t = 0.2, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 15 97
4.59 Model 1 (a/t = 0.2, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 10 97
4.60 Model 1 (a/t = 0.2, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 5 98
4.61 Model 3 (a/t = 0.2, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 15 98
4.62 Model 3 (a/t = 0.2, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 10 99
4.63 Model 3 (a/t = 0.2, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 5 99
4.64 Model 4 (a/t = 0.5, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 15 100
4.65 Model 4 (a/t = 0.5, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 10 100
-
xviii
Figure Page
4.66 Model 4 (a/t = 0.5, a/c = 0.2): h1 vs. c/w for Nasgro and FEM at n = 5 101
4.67 Model 6 (a/t = 0.5, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 15 101
4.68 Model 6 (a/t = 0.5, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 10 102
4.69 Model 6 (a/t = 0.5, a/c = 1.0): h1 vs. c/w for Nasgro and FEM at n = 5 102
B.1 Model 1 (a/t=0.2, a/c=0.2): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 116
B.2 Model 2 (a/t=0.2, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 116
B.3 Model 3 (a/t=0.2, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117
B.4 Model 4 (a/t=0.5, a/c=0.2): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 117
B.5 Model 5 (a/t=0.5, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118
B.6 Model 6 (a/t=0.5, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118
B.7 Model 8 (a/t=0.8, a/c=0.6): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119
B.8 Model 9 (a/t=0.8, a/c=1.0): Normalized K factor vs. angle along crackfront . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119
C.9 Model 1: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 121
C.10 Model 1: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 121
C.11 Model 1: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 122
C.12 Model 2: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 122
-
xix
Figure Page
C.13 Model 2: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 123
C.14 Model 2: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 123
C.15 Model 3: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 124
C.16 Model 3: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 124
C.17 Model 3: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 125
C.18 Model 4: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 125
C.19 Model 4: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 126
C.20 Model 4: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 126
C.21 Model 5: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 127
C.22 Model 5: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 127
C.23 Model 5: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 128
C.24 Model 6: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 128
C.25 Model 6: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 129
C.26 Model 6: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 129
C.27 Model 8: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 130
C.28 Model 8: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 130
C.29 Model 8: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 131
C.30 Model 9: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 131
C.31 Model 9: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 132
C.32 Model 9: h1 vs. angle along the crack front . . . . . . . . . . . . . . . . 132
-
LIST OF SYMBOLS
Symbol Description
a Crack depthaeff Effective crack length, includes plastic zoneb Uncracked ligament lengthc Half crack lengthds Increment of length along the contourG Strain energy release rateh1 Dimensionless parameter used to calculate Jplh2 Dimensionless parameter used to calculate CTODh3 Dimensionless parameter used to calculate δpn Strain hardening exponentnj Unit vector components normal to Γr Crack tip radiusrc Radius of projected circlet Specimen thicknessw Half specimen widthx1 Distance along x-axis for projected circlex2 Distance along x-axis for projected circley1 Distance along y-axis for projected circley2 Distance along y-axis for projected circleA Crack areaCTOD Crack tip opening displacementE Young’s ModulusIn Integration constantJ Elastic-plastic fracture parameterJel Elastic portion of the J-integralJpl Plastic portion of the J-integralJtotal Sum of Jel and JplK Stress intensity factorKnorm Normalized K-factorP Applied loadPo Limit loadTi Traction vectorui Displacement vectorW Specimen widthα Dimensionless Ramberg-Osgood material constantβ Plasticity constraint factorδp Load line displacement
xx
-
xxi
Symbol Description
� Strainεij Strain tensor�o Yield strain�ref Reference strainΓ ContourΠ Potential Energyµ Reference stress factorν Poisson’s ratioω Strain energy densityσ Stressσij Stress tensorσo Yield stressσref Reference stressθ Angle of crack tipEPFM Elastic plastic fracture mechanicsEPRI Electric Power Research InstituteFEA Finite element analysisFEM Finite element modelLEFM Linear elastic fracture mechanicsODB Output data base
-
CHAPTER 1
INTRODUCTION
1.1 Fracture Mechanics
Fracture mechanics is the study of the effects of flaws in materials under load.
Modern fracture mechanics was originated by Griffith [1] in the 1920’s when he suc-
cessfully showed that fracture in glass occurs when the strain energy resulting from
crack growth is greater than the surface energy. In 1948, Irwin [2] extended Griffith’s
strain energy release rate, G, to include metals by accounting for the energy absorbed
during plastic material flow around the flaw. By 1960, the fundamental principles of
linear elastic fracture mechanics (LEFM) were in place ([3, 4], for example).
LEFM is used to predict material failure when response to the load is elastic
and the fracture response is brittle. LEFM uses the strain energy release rate G or
the stress intensity factor K as a fracture criterion. K solutions for many geometries
have been calculated in the past and are widely available [5]. However, the design
parameters for some components violate the assumptions of LEFM. For example,
high temperatures and limited high stress cycles before component replacement are
factors that can cause significant plastic deformation and a ductile failure. In these
cases, where the LEFM approach is not valid, an elastic-plastic fracture mechanics
(EPFM) approach is required.
EPFM had its beginnings in 1961, when Wells [6] noticed that initially sharp
cracks in high toughness materials were blunted by plastic deformation. Wells pro-
posed that the distance between the crack faces at the deformed tip be used as a
1
-
2
measure of fracture toughness. The stretch between the crack faces at the blunted
tip is known as the crack tip opening displacement (CTOD).
In 1968 Rice [7] developed another EPFM parameter called the J-integral by
idealizing the elastic-plastic deformation around the crack tip to be nonlinear elastic.
The J-integral was shown to be equivalent to G for linear elastic deformation and to
the crack tip opening displacement for elastic-plastic deformation. During the same
year, Hutchinson [8], Rice, and Rosengren [9] showed that J was also a nonlinear
stress intensity parameter. The J-integral can be used as an elastic-plastic or fully
plastic crack growth fracture parameter, much like K is used as an elastic fracture
parameter.
The J-integral can be calculated using several experimental and analytical
techniques. The analytical techniques include the Electric Power Research Institute
(EPRI) estimation scheme, the reference stress method, and finite element methods.
It should be noted that many of the analytical techniques that do not directly require
finite element methods were established using finite element analysis.
1.2 Overview of Research
There are three goals in this research. The primary goal is to develop three-
dimensional finite element analysis (FEA) J-integral results using ABAQUS. These
results will be compared to existing solutions. The second goal is to investigate
the effect of various finite element modelling parameters on the resulting J-integral.
These parameters include mesh density, element type, symmetry, and specimen size
effects. The third goal is to compare incremental plasticity FEAs that utilize a stress
vs. plastic strain table based on a power law hardening material with the deformation
plasticity solution for a power law material. This comparison will be made in an
-
3
attempt to see if the fully plastic results using a deformation plasticity model can be
approached by a series of increasing loads in an incremental plasticity model.
The finite element models (FEMs) used in this research were three-dimensional
flat plates with surface cracks. The plates contained various surface crack, height, and
width geometries. Because of the dual symmetry, only one quarter of each plate was
modeled. Meshes from two different mesh generation programs were used: mesh 3d
(Faleskog, 1996) and FEA-Crack from Structural Reliability Technology.
-
CHAPTER 2
TECHNICAL BACKGROUND
In this chapter the J-integral and different J-integral calculation methods will
be examined. The chapter begins with a discussion of the theory and mathematical
foundation of the J-integral. Next, two methods for calculating the J-integral are
discussed: the EPRI Estimation Scheme and the reference stress method. Both of
these methods can be implemented using “hand calculations” without an extensive
fracture mechanics background. In addition, both of these methods are incorporated
into Nasgro, a fracture mechanics and fatigue crack growth program. Finally, the
FEA method is used in this research, but a review is not included here. There are
many excellent texts on the subject of FEA (for example Cook et al. [10]).
2.1 J-Integral
Rice [7] developed J as a path-independent contour integral by idealizing
elastic-plastic deformation to be the same as nonlinear elastic material behavior. In
the arbitrary path around a crack tip (Figure 2.1),
J =
∫Γ
(ωdy − Ti
∂ui∂x
ds
), (2.1)
where ω is the strain energy density, Ti are components of the traction vector, ui
are the displacement vector components, and ds is an increment of length along the
4
-
5
Figure 2.1 Contour around a crack tip [4]
contour(Γ). The strain energy density and the traction vector components are
ω =
εij∫0
σijdεij (2.2)
and
Ti = σijnj, (2.3)
where σij is the stress tensor, εij is the strain tensor, and nj are unit vector components
normal to Γ.
In idealizing elastic-plastic behavior to be the same as nonlinear elastic material
behavior, Rice assumed that the material stress versus strain curve followed a power
law relationship. The Ramberg-Osgood equation is commonly used to describe the
stress and total strain data for this type of material response:
ε
εo=
σ
σo+ α
(σ
σo
)n, (2.4)
-
6
where ε is the total material strain, σo is the reference stress (normally defined as the
yield strength, but not necessarily the same as the 0.2% offset yield strength), εo is
the strain at the reference stress and is defined by εo = σo/E. There are two other
material constants in Equation 2.4. The first of these, α, is a dimensionless constant,
and the second, n, is the strain hardening exponent (n ≥ 1).
The J-dominated elastic-plastic stress field contains a singularity of order
r−1
n+1 . For the elastic case (n = 1), this singularity reduces to r−12 in agreement
with the K-dominated field of LEFM. The following two equations were derived by
Hutchinson [8], Rice and Rosengren [9] and are called the HRR singularity. The HRR
singularity describes the actual stresses and strains near the crack tip and within the
plastic zone as
σij = σo
(EJ
ασ2oInr
) 1n+1 ∗
σij (n, θ) (2.5)
and
εij =ασoE
(EJ
ασ2oInr
) nn+1 ∗
εij (n, θ) , (2.6)
where In is an integration constant depending on n, r is the crack tip radius, θ is
the angle at a point around the contour, and∗
σij and∗
εij are functions of n and θ.
Equations 2.5 and 2.6 are important because the J-integral determines the stress
amplitude within the plastic zone. This fact establishes J as a fracture parameter
under conditions of plastic deformation.
-
7
Rice [7] also showed that the J-integral is equivalent to the energy release rate
in a nonlinear elastic material containing a crack:
J = −dΠdA
(2.7)
where Π is the potential energy and A is the area of the crack. For linear elastic
deformation:
Jel = G =K2
E ′(2.8)
where, for plane strain
E ′ =E
(1− ν2), (2.9)
and, for plane stress
E ′ = E. (2.10)
Care should be taken when using the energy release rate with elastic-plastic
or fully plastic deformation. In an elastic material, the potential energy is released
as the crack grows. In an elastic-plastic material, a large amount of strain energy is
used in forming a plastically deformed region around the crack tip. This energy will
not be recovered when the crack grows, or when the specimen is unloaded [4].
-
8
2.2 EPRI Estimation Scheme
The elastic-plastic and fully plastic J-integral estimation scheme presented by
EPRI [11] is derived from the work of Shih [12] and Hutchinson [13]. The purpose
of this work was to devise a simple handbook-style procedure for calculating the J-
integral. This goal was made possible by compiling nondimensional functions in table
form that could be used to calculate J directly. The nondimensional functions were
based on FEA results using Ramberg-Osgood materials.
The EPRI procedure computes a total J by summing the elastic and plastic
J ’s for various 2D geometries. This is expressed as
Jtotal = Jel + Jpl (2.11)
where Jtotal is the total J , Jel is the elastic portion, and Jpl is the plastic portion. For
small loads, Jel is much larger than Jpl. For large loads with significant deformation,
Jpl dominates. This situation is shown graphically in Figure 2.2. As discussed previ-
ously, elastic-plastic behavior is idealized to follow a nonlinear elastic path along the
stress versus strain curve.
In the EPRI estimation scheme, Jel is calculated utilizing an adjusted crack
length (aeff ) to compensate for the strain hardening around the crack tip and is
expressed as
Jel = G =K2(aeff )
E ′, (2.12)
-
9
Figure 2.2 EPRI J-Integral estimation scheme [4]
where K is the stress intensity factor as a function of aeff . The adjusted crack length
is given by
aeff = a +1
1 + (P/Po)2
1
βπ
(n− 1n + 1
) (KIσo
)2, (2.13)
where a is the half crack length, P is the applied load, Po is the limit load per unit
thickness, β = 2 for plane stress and β = 6 for plane strain, n is the strain hardening
exponent specific to the material, KI is the elastic stress intensity factor, and σo is
the reference stress (typically the yield strength).
-
10
The fully plastic equations for Jpl, crack mouth opening displacement (CTOD),
and load line displacement (δp), applicable for most specimen geometries are
Jpl = αεoσobh1
( aW
, n) ( P
Po
)n+1, (2.14)
CTOD = αεoah2
( aW
, n) ( P
Po
)n, (2.15)
and
δp = αεoah3
( aW
, n) ( P
Po
)n, (2.16)
where α and n are a material constants, b is the uncracked ligament length, W is
the specimen width, and a is the crack length. h1, h2, and h3 are dimensionless
parameters that are a function of geometry and the hardening exponent n.
The center-cracked and single-edge-notched specimen geometries have a dif-
ferent form for Jpl. This form reduces the effect of the crack length to width ratio on
the value of h1, and is
Jpl = αεoσoba
wh1
( aw
, n) ( P
Po
)n+1, (2.17)
where, for a center-cracked specimen, a is the half crack length and w is the half
width. Po is the reference or limit load, and is typically the load at which net cross
section yielding occurs. For center-cracked plate in tension,
Po = 4cσo
/√3 for plane strain, (2.18)
-
11
and
Po = 2cσo for plane stress. (2.19)
For a single-edge-crack in tension,
Po = 1.455ηcσo for plane strain, (2.20)
and
Po = 1.072ηcσo for plane stress. (2.21)
The EPRI handbook includes tabulations of h1, h2, and h3 for various n values
and geometries. These values were calculated using results from a finite element pro-
gram called INFEM [11]. INFEM was developed for the specific purpose of analyzing
fully plastic cracks and utilizes incompressible elements in the model formulation.
Further details of the finite element formulation have been published by Needleman
and Shih [14].
In 1999 McClung, Chell, Lee, and Orient [15] extended the original EPRI work
to include fully plastic J solutions for 3D geometries. This work was performed using
3D finite element models. The meshes for these models were constructed using eight-
noded brick elements in ANSYS 5.0. A typical mesh is shown in Figure 2.3. A close
up view of the crack front may be seen in Figure 2.4.
-
12
Figure 2.3 Sample of finite element mesh used by McClung et al. [15]
Figure 2.4 Close up of the finite element mesh around the crack front used byMcClung et al. [15]
-
13
Although the meshes were created in ANSYS, ABAQUS was used to perform
the analysis of the finite element models. The version of ABAQUS used for this work
was only capable of performing an incremental plasticity analysis. An EPRI-type
scheme was used to separate the elastic and plastic J values. The fully plastic values
for h1 were then calculated using
h1 =Jpl
ασoεot(
σσo
)n+1 . (2.22)A combination of three different a/t (0.2, 0.5, 0.8) and a/c (0.2, 0.6, and 1.0)
ratios were tabulated. The specimen geometry ratios were kept constant for all models
at h/c = 4 and c/w = 0.25. The values of h1 were calculated for strain hardening
exponents of n = 5, 10, and 15, and can be found in Tables 2.1, 2.2, and 2.3.
In 2004 Lei [17] duplicated part of the work performed by McClung et al. [15]
by performing elastic and elastic-plastic finite element analyses for plates containing
semi-elliptical surface cracks under tension. The models contained surface cracks with
the same a/t and a/c ratios used by McClung et al. [15]. For the elastic analysis,
Jel results were generated and converted into K using Equation 2.8. These K results
were then compared with Newman-Raju stress-intensity factor calculations [18]. The
elastic-plastic results for strain hardening values of n = 5 and n = 10 were presented
in terms of h1. These h1 results are reproduced in Tables 2.4 and 2.5 and compare
well with McClung et al. for most geometries. The comparison with McClung et
al. and the current results are presented in more detail in Chapter 4.
-
14
Tab
le2.
1M
cClu
ng
etal
.h
1va
lues
inte
nsi
on,n
=15
[15]
a/t
a/c
0◦
9◦
18◦
27◦
36◦
45◦
54◦
63◦
72◦
81◦
90◦
0.20
0.20
0.22
30.
370
0.60
80.
821
1.00
11.
148
1.31
01.
447
1.56
01.
623
1.64
40.
200.
600.
356
0.46
50.
622
0.69
80.
774
0.82
30.
875
0.91
50.
948
0.97
10.
981
0.20
1.00
0.38
90.
503
0.62
80.
638
0.65
90.
653
0.65
70.
657
0.65
30.
646
0.64
60.
500.
204.
085
7.61
511
.602
14.4
8817
.057
18.7
9820
.228
21.4
3422
.129
22.2
1222
.309
0.50
0.60
3.33
64.
808
6.56
47.
048
7.69
77.
939
8.00
08.
021
8.01
97.
922
7.88
10.
501.
002.
774
3.73
84.
750
4.75
94.
932
4.89
14.
816
4.61
34.
407
4.24
34.
198
0.80
0.20
37.6
0963
.511
82.4
0491
.460
99.1
9892
.725
90.0
9788
.292
89.5
4895
.447
98.9
410.
800.
6017
.660
25.8
9032
.760
30.1
7237
.828
34.0
0231
.546
28.9
7228
.224
29.0
9530
.806
0.80
1.00
12.6
6717
.231
20.8
8219
.281
23.0
2921
.124
18.0
0516
.467
14.5
5715
.003
15.5
33
Tab
le2.
2M
cClu
ng
etal
.h
1va
lues
inte
nsi
on,n
=10
[15]
a/t
a/c
0◦
9◦
18◦
27◦
36◦
45◦
54◦
63◦
72◦
81◦
90◦
0.20
0.20
0.19
80.
320
0.52
30.
703
0.86
30.
996
1.13
31.
250
1.34
51.
398
1.41
60.
200.
600.
324
0.41
60.
544
0.60
40.
671
0.71
50.
759
0.79
20.
820
0.83
90.
847
0.20
1.00
0.35
80.
450
0.55
00.
553
0.57
10.
565
0.56
90.
566
0.56
20.
557
0.55
60.
500.
202.
539
4.51
26.
957
8.84
110
.665
11.9
7913
.048
13.9
5314
.546
14.7
1214
.811
0.50
0.60
2.31
93.
205
4.26
44.
561
4.96
75.
128
5.18
95.
209
5.23
15.
200
5.18
60.
501.
002.
007
2.59
93.
210
3.17
93.
272
3.21
83.
168
3.04
02.
921
2.82
72.
804
0.80
0.20
17.7
3129
.512
39.5
5043
.774
49.6
0046
.576
44.8
5443
.844
43.7
0645
.805
47.4
960.
800.
609.
688
13.6
8516
.725
15.8
5019
.174
17.3
1815
.896
14.7
7614
.068
14.3
2314
.800
0.80
1.00
7.24
29.
472
11.0
7710
.311
11.8
9811
.108
9.23
98.
398
7.53
67.
533
7.62
5
-
15
Tab
le2.
3M
cClu
ng
etal
.h
1va
lues
inte
nsi
on,n
=5
[15]
a/t
a/c
0◦
9◦
18◦
27◦
36◦
45◦
54◦
63◦
72◦
81◦
90◦
0.20
0.20
0.16
40.
252
0.40
70.
544
0.67
60.
789
0.89
70.
988
1.06
21.
103
1.11
70.
200.
600.
286
0.35
20.
441
0.48
00.
533
0.57
00.
605
0.63
10.
652
0.66
60.
672
0.20
1.00
0.32
10.
383
0.44
60.
440
0.45
20.
446
0.44
70.
442
0.43
90.
435
0.43
50.
500.
201.
325
2.13
93.
357
4.38
45.
480
6.37
17.
136
7.76
48.
222
8.42
88.
516
0.50
0.60
1.50
21.
916
2.41
22.
548
2.76
42.
860
2.91
72.
938
2.96
82.
973
2.97
60.
501.
001.
377
1.65
81.
931
1.86
71.
894
1.83
91.
800
1.73
01.
677
1.63
71.
630
0.80
0.20
7.22
411
.273
15.7
4318
.150
21.8
7021
.460
20.6
3220
.051
18.9
6018
.993
19.3
690.
800.
604.
983
6.44
97.
582
7.38
98.
421
7.75
07.
034
6.69
56.
266
6.11
46.
178
0.80
1.00
3.91
04.
728
5.25
14.
849
5.14
24.
775
4.08
03.
734
3.39
73.
285
3.27
0
-
16
Tab
le2.
4Lei
h1
valu
esin
tensi
on,n
=5
[17]
a/t
a/c
0◦
9◦
18◦
27◦
36◦
45◦
54◦
63◦
72◦
81◦
90◦
0.2
0.2
0.17
90.
3003
0.45
720.
5949
0.72
230.
8389
0.95
561.
053
1.13
21.
177
1.19
60.
20.
60.
3151
0.39
580.
4897
0.52
360.
5777
0.61
340.
6517
0.67
780.
7007
0.71
130.
7177
0.2
10.
3575
0.43
680.
5011
0.48
70.
5004
0.48
950.
4912
0.48
580.
4863
0.48
250.
4839
0.5
0.2
1.34
32.
327
3.49
54.
524
5.45
56.
265
7.03
97.
616
8.08
98.
338
8.46
60.
50.
61.
564
2.03
2.52
42.
633
2.82
92.
897
2.97
42.
982.
993
2.97
22.
981
0.5
11.
441.
782.
042
1.95
1.97
1.87
81.
837
1.76
21.
722
1.67
61.
672
0.8
0.2
6.72
311
.78
17.2
519
.57
21.2
21.4
321
.16
20.3
519
.57
18.9
218
.86
0.8
0.6
5.38
86.
938
8.31
78.
145
8.18
57.
656
7.17
86.
633
6.39
66.
263
6.29
0.8
14.
119
5.07
55.
678
5.18
75.
014
4.48
24.
106
3.70
13.
505
3.39
63.
406
Tab
le2.
5Lei
h1
valu
esin
tensi
on,n
=10
[17]
a/t
a/c
0◦
9◦
18◦
27◦
36◦
45◦
54◦
63◦
72◦
81◦
90◦
0.2
0.2
0.21
690.
3749
0.57
140.
7474
0.90
661.
046
1.18
61.
302
1.39
71.
451
1.47
50.
20.
60.
3554
0.45
080.
5774
0.63
30.
7033
0.75
180.
7987
0.83
380.
8627
0.87
860.
8862
0.2
10.
3995
0.50
090.
5964
0.59
950.
6222
0.61
780.
6211
0.61
90.
6193
0.61
720.
618
0.5
0.2
2.53
34.
723
6.90
78.
859
10.3
411
.45
12.4
513
.09
13.6
613
.91
14.1
10.
50.
62.
379
3.25
44.
285
4.58
84.
954
5.07
15.
164
5.14
25.
122
5.07
5.07
70.
51
2.05
52.
682
3.26
3.25
33.
341
3.23
13.
148
3.01
22.
907
2.81
82.
799
0.8
0.2
16.0
129
.51
43.2
746
.42
47.1
745
.94
45.7
145
.34
45.1
144
.93
45.3
0.8
0.6
11.0
815
.38
19.3
19.1
319
.01
17.4
316
.15
15.2
315
.31
15.4
815
.67
0.8
17.
925
10.6
812
.74
12.1
311
.77
10.4
59.
369
8.43
98.
159
8.23
8.40
8
-
17
2.3 Reference Stress Method
As discussed previously, the EPRI J estimation scheme assumes that the mate-
rial has a power law stress-strain curve. There are many materials that do not exhibit
this type of response. In 1984 Ainsworth [19] devised a method for calculating J that
did not depend on the material’s behavior following a power law. This approach is
called the reference stress method. The reference stress is defined as
σref =
(P
Po
)σo (2.23)
where P is the applied load, Po is the same limit load defined previously in the EPRI
research [11], and σo is the yield strength.
The reference strain, εref , is defined as the uniaxial strain corresponding to
σref . By inserting σref and εref into the Ramberg-Osgood equation 2.4, it can be
modified to the following form:
εrefεo
=σrefσo
+ α
(σrefσo
)n. (2.24)
Using Equations 2.23 and 2.24, Equation 2.14 can be altered to the form
Jpl = σrefbh1
(εref −
σrefεoσo
). (2.25)
Equation 2.25 still contains the variable h1, a function of n - same h1 used in the EPRI
equations discussed in the previous section. Ainsworth’s approach was to choose Po
in such a way that the dependence of h1 on n was minimized. For certain values of
-
18
Po, he found that h1 was relatively constant for n ≤20. As a result,
h1 ∼= h1( a
w, 1
)(2.26)
where h1 is the average h1 for a range of n’s and h1(
aw, 1
)is the h1 for n equal to one.
The fully plastic solution at n = 1 is identical to the elastic solution using a Poisson’s
ratio of υ = 0.5,
µK2 (a) = bh1
( aw
, 1)
σ2ref (2.27)
where µ=1 for plane stress and µ=0.75 for plane strain. By substituting Equation
2.27 and using the conditions that establish Equation 2.26, the Jpl expression becomes
Jpl =µKIE
(Eεrefσref
− 1)
. (2.28)
The previously discussed McClung et al. [15] finite element results were used
to develop another reference stress method. This reference stress algorithm is used
within Nasgro. Nasgro is a crack propagation and fracture mechanics program devel-
oped by NASA and the Southwest Research Institute.
-
CHAPTER 3
RESEARCH PROCEDURE
In this chapter, the technical approach used for this thesis is presented. The
chapter begins with a discussion of the finite element modeling including mesh gen-
eration. Next, the analysis procedure for the FEMs is discussed. Then, the work
duplicated by other researchers is reviewed, and any material properties or model
parameters specific to a geometry set are looked at as well. This duplication of other
researchers’ work was to validate the methodology used by ensuring that the J-integral
analysis could be performed properly. The chapter concludes with a discussion of the
general material properties used.
3.1 Finite Element Modeling
The finite element analysis program ABAQUS was used to calculate the K-
factors and J-integrals for a variety of specimen geometries. The models were created
with quarter symmetry to reduce the number of nodes and elements (hence, the
computational time) of each model.
Unless otherwise specified, the FEMs consisted of reduced integration, 20-
noded brick elements specified as C3D20R within ABAQUS. Reduced integration
elements are recommended in the ABAQUS User Manuals [21] for plastic and large
strain elastic models. Full integration elements tend to be overly stiff and the results
may oscillate. A reduced integration element has a softening effect on the stiffness
that improves the finite element results.
The elements around the crack tip were also of type C3D20R. However, the
elements were modified by collapsing the brick element into a wedge (Figure 3.1).
19
-
20
When the elements were degenerated, the mid-side nodes were not moved, and the
collapsed nodes were left untied (Figure 3.2). This allows for movement of the nodes
as the element is deformed and produces a 1/r strain singularity, which duplicates
the actual crack tip strain field in the plastic zone [4].
Figure 3.1 Degeneration of elements around crack tip [4]
Figure 3.2 Plastic singularity element [4]
-
21
3.1.1 Mesh Generation
Two different programs were used to generate finite element meshes. The
first, called mesh3d scp [20] by Faleskog, is available as freeware. Many early finite
element meshes in this work were generated with mesh3d scp. However, this program
has serious limitations. Therefore, a second mesh generation program, FEA-Crack,
was also used. This software is commercially available from Structural Reliability
Technology, Colorado.
3.1.1.1 mesh3d scp. The mesh generation program mesh3d scp generates
a one-quarter model of a surface cracked plate. The program assumes that both the
geometry and the load possess planes of symmetry. This program divides the model
into three zones, as shown in Figure 3.3. The element density in each zone is altered by
changing variables in the mesh3d scp input file. The node and element numbering in
each zone is controlled such that the application of boundary conditions and external
loads is simplified. The meshes used to investigate the fully plastic volume and
location were created using mesh3d scp (Figures 3.8 - 3.11).
The program mesh3d scp requires an iterative approach. The set of input
variables for the program input file are changed, the program generates a mesh, the
mesh is plotted and then examined graphically. This process is repeated until a
satisfactory mesh by appearance is created. This program is capable of generating
good meshes for some geometries. However, this program does not work well for other
specimen geometries. For these geometries, mesh3d scp was found to produce a bad
mesh, no mesh, or, in the worst cases, a mesh with errors.
This program was originally written to generate meshes for an earlier version
of ABAQUS. This makes it necessary to modify the ABAQUS input files created by
-
22
Figure 3.3 Zones created in the mesh by mesh3d scp [20]
mesh3d scp to make them compatible with recent releases of ABAQUS(V6.5). The
file modifications used for the models in this thesis are listed in Appendix A.
3.1.1.2 FEA-Crack. The second mesh generation program utilized for this
research is called FEA-Crack. FEA-Crack is more robust than mesh3d scp and does
not require the same iterative approach on the user’s part. The mesh density in the
area around the crack can be controlled by adjusting the program settings. Also, the
generated model may be viewed immediately, and required changes to the ABAQUS
input file are minimal. A mesh created using FEA-Crack is shown in Figures 3.4 and
3.5.
-
23
Figure 3.4 Mesh created using FEA-Crack
Figure 3.5 Close up of mesh from Figure 3.4 created using FEA-Crack
-
24
3.2 Analysis Procedure
Each FEM analyzed for this research contained 5 contours around the crack
tip, as seen in Figure 3.6. The results for the first contour are generally considered
to be less accurate than the other contours because of numerical inaccuracy [21]. For
this reason, the K-factor and J-integral data from all of the contours, except the first,
were averaged [17]. These average K-factor and J-integral were used for all further
calculations and comparisons.
The FEMs contained multiple node sets along the crack front. A node set
is a group of nodes that have been associated as a group within ABAQUS. The
number of node sets depended on the physical size of the crack front. Each of these
particular node sets contain a number of nodes with the same coordinates. In the
untied condition, one node in each node set is constrained so that it can move in only
Figure 3.6 Contours (semi-circular rings) around the crack tip
-
25
one or two directions (it stays on the plane of symmetry). The direction of constraint
depends on the symmetry plane. These constrained nodes are listed in another node
set called “crack front nodes,” which will be significant later. The other nodes in
each node set are not constrained.
ABAQUS generates values for the K-factor and J-integral at each of the node
sets along the crack front. An Excel macro was written to allow for examination of
the variation of the K-factor and J-integral values generated along the crack front.
The program was written to calculate the angle, as projected onto a circle, at each
crack front node. The macro first finds and records the constrained nodes found in
the node set “crack front nodes,” which is located in the ABAQUS input file. The
coordinates for each of these crack front nodes are then retrieved from the input file.
The crack coordinates are then mapped onto a circle, as shown in Figure 3.7. The
equation for the projection circle is shown below as
x22 + y22 = r
2c . (3.1)
Two facts should be noted from Figure 3.7. First, y1 is equal to y2. Second, the
circle radius, rc, is equal to the crack depth, a. Both of the previous statements are
valid as long as a/c ≤ 1, which is the case for this research. Using this information,
Equation 3.1 can now be rearranged into the form
x2 =√
a2 − y21. (3.2)
Once x2 is known, the angle, θ, may be calculated using
θ = tan−1(
x2y1
). (3.3)
-
26
Figure 3.7 Coordinate scheme for mapping crack face angles
With θ known, the variation of the K-factor and J-integral values can be mapped
along the crack front contour.
3.3 J-Integral Convergence
Two quantities were initially tested to ensure that the fully plastic FEM results
had converged. The first quantity was load. The second involved the fully plastic
zone specified for the FEMs.
3.3.1 Load
The applied load in the FEMs was adjusted until the resulting J-integral values
did not change with an increase in load. The final load step was also examined for
each model to ensure that the entire load was not applied. In cases where the entire
-
27
specified load was applied, the load was increased, and the FEM was analyzed again.
This ensured that the specified element set became fully plastic. The fully plastic
option in ABAQUS utilizes a Ramberg-Osgood material model and ends the analysis
when the observed strain for the selected element set exceeds the offset yield strain
by ten times, assuming the load or maximum number of increments have not been
reached. Also, to ensure sufficient steps in the model, the loads were set such that at
least 33% of the specified load was applied to the model.
3.3.2 Fully Plastic Zone
The volume and location effect of the specified fully plastic element set was
examined for two reasons. First, it was necessary to determine how much of the
specimen must become fully plastic before the J-integral converged. The second
reason was to simplify the model generation. The two mesh generation programs used
in this research, mesh3d scp and FEA-Crack, established convenient, but different,
elements sets for use as fully plastic.
The fully plastic results were generated using the *FULLY PLASTIC command
within ABAQUS. This command requires the specification of an element set which
is monitored for the fully plastic condition discussed previously. Several fully plastic
element sets, or zones, were tested and the results compared. The fully plastic element
sets used in this research are defined as follows:
• LayerCR - Contains elements around the crack tip, (Figure 3.8);
-
28
Figure 3.8 Fully plastic element set consisting of the elements around the cracktip
• Partial Layer 1 - Contains elements in the first layer of the model, but does
not contain the elements closest to the crack tip, (Figure 3.9);
• Layer 1 - Contains the elements in the ligament plus the elements found in
LayerCR, (Figure 3.10);
• Layer 2 - Contains elements in the first and second layers of the model, but
does not contain the elements closest to the crack tip (Figure 3.11).
-
29
Figure 3.9 Fully plastic element set consisting of part of layer 1
Figure 3.10 Fully plastic element set consisting of layer 1
-
30
Figure 3.11 Fully plastic element set consisting of partial layers 1 and 2
-
31
3.4 Comparison to Other Work
A series of models with different crack ratios and specimen sizes were generated.
These models contained geometric parameters (e. g. a/t, a/c, etc.) identical to those
used by other researchers. The current results were compared to previous work with
the intent of validating the FEMs and methods used for this research.
3.4.1 Kirk and Dodds
FEMs were generated with the same geometries and material properties used
by Kirk and Dodds in 1992 [23]. These geometries are shown in Figure 3.12. The
mesh generation program mesh3d scp was used to generate models for all three cracks
defined by Kirk and Dodds. The models consisted of 20-noded brick elements with
reduced integration. The number of nodes and elements in each model is listed in
Table 3.1.
Table 3.1 Number of nodes and elements in the duplication of the Kirk and Dodds[23] geometries
Crack 1 Crack 2 Crack 3Nodes 16,597 12,227 12,227
Elements 3562 2593 2593
-
32
Figure 3.12 Geometries used by Kirk and Dodds for estimating the J-Integral [23]
-
33
These FEMs were analyzed to find Jtotal using an elastic-plastic analysis.
ABAQUS utilizes an incremental plasticity model for this type of analysis, and re-
quires a table of true stress versus plastic strain. The material properties for these
models were derived from Figure 3.13 and are listed below:
• E = 3.00× 104 kpsi
• ν = 0.3
• Tangent Modulus = 3.57× 102 kpsi
• Initial Yield = 80 kpsi.
These properties were used to calculate the total and elastic strains at the yield stress
and an arbitrary stress, selected to be much higher than the applied stress. This
arbitrarily large stress was used as an input because ABAQUS does not explicitly
allow the tangent modulus to be given. The plastic strains required by ABAQUS
were found by subtracting the total and elastic strains. Table 3.2 shows the calculated
strains.
Table 3.2 Incremental plasticity values for the Kirk and Dodds models
σ, kpsi total strain elastic strain plastic strain80 2.67E-03 2.67E-03 0.00E+00200 3.36E-01 6.67E-03 3.29E-01
-
34
Figure 3.13 Stress vs. strain curve for Kirk and Dodds elastic-plastic models [23]
-
35
3.4.2 McClung et al. [15]
The mesh generation program FEA-Crack was used to generate models for all
nine geometries defined in the research performed by McClung et al. (Table 3.3). Two
sets of models were generated. The first set contained a coarse mesh. The second set
utilized a more refined mesh around the crack front. The McClung et al. geometries
were analyzed as elastic, fully plastic and incrementally plastic models. The elastic
and fully plastic analyses were performed using both the coarse and refined meshes.
The incrementally plastic models were analyzed using only the coarse meshes.
In the elastic FEM analysis, the K factor was found in two ways. First,
ABAQUS was used to calculate K directly. Second, ABAQUS was used to find
the elastic J , and then Equation 2.8 was used to calculate K. These results were
compared to K factors calculated using equations from Newman and Raju [24]. The
Newman-Raju solution is given in Equations 3.4 - 3.9.
KI = σ
√π
(a
Q
) [M1 + M2
(at
)2+ M3
(at
)4]gfθfw, (3.4)
Q = 1 + 1.464(a
c
)1.65, (3.5)
-
36
Tab
le3.
3M
cClu
ng
etal
.fu
lly
pla
stic
geom
etries
Model1
Model2
Model3
Model4
Model5
Model6
Model7
Model8
Model9
a/t
0.2
0.2
0.2
0.5
0.5
0.5
0.8
0.8
0.8
a/c
0.2
0.6
10.
20.
61
0.2
0.6
1h/c
44
44
44
44
4c/
w0.
250.
250.
250.
250.
250.
250.
250.
250.
25t
11
11
11
11
1a
0.2
0.2
0.2
0.5
0.5
0.5
0.8
0.8
0.8
c1
0.33
0.2
2.5
0.83
0.5
41.
330.
8w
41.
330.
810
3.33
216
5.33
3.2
h4
1.33
0.8
103.
332
165.
333.
2
-
37
M1 = 1.13− 0.09(
ac
),
M2 = −0.54 + 0.890.2+(ac ),
M3 = 0.5− 10.65+ac
+ 14(1− a
c
)24,
(3.6)
g = 1 +
[0.1 + 0.35
(at
)2](1− sin θ)2 , (3.7)
fθ =
[(ac
)2cos2 θ + sin2 θ
]1/4, (3.8)
fw =
[sec
(πc
2w
√a
t
)]1/2, (3.9)
where KI is the K factor at a given angle, σ is the applied stress, a is the crack depth,
Q is factor applicable for ac≤ 1, c is the half crack width, t is the specimen thickness,
θ is the angle, as previously defined in Figure 3.7, along the crack front, and w is the
half specimen width.
3.4.3 Lei [17]
In 2004, Lei performed elastic and elastic-plastic J analyses on models with
the same crack geometries used by McClung et al. [15]. He also maintained a spec-
imen geometry ratio of c/w = 0.25. However, Lei deviated from the McClung et
-
38
al. geometries by fixing the ratio h/w at four to one instead of one to one. Lei also
fixed c, therefore fixing w and h, and varied a and t.
Lei used ABAQUS to perform the analyses on his models. He used the *CON-
TOUR INTEGRAL command within ABAQUS to generate J-integral results for
fifteen contours around the crack tip. The averages of these contours, excluding the
first, were presented. Lei found that the deviation of data from any one contour is
less than 5% of the average value.
Lei used consistent material properties in his analyses. The properties for the
elastic analyses were set at E = 500 MPa and ν = 0.3. The elastic-plastic analyses
used the Ramberg-Osgood stress-strain relationship (Equation 2.4), where σo = 1.0
MPa, α = 1, and n = 5 and 10. For all analyses, Lei used the Mises yield criterion
and small strain isotropic hardening.
3.4.4 Nasgro Computer Program
Current FEM results were compared with the results produced using the crack
propagation and fracture mechanics section of Nasgro. Nasgro is a fracture mechanics
and fatigue crack growth program developed by NASA and the Southwest Research
Institue. The same Ramberg-Osgood material properties used for the McClung ge-
ometries were duplicated for this comparison. The different geometries analyzed using
Nasgro are shown in Table 3.4.
-
39
Table 3.4 Geometries for Nasgro comparison and width effect investigation
Model a a/t c c/w w1 0.2 0.2 1.0 0.25 4.001a 0.2 0.2 1.0 0.50 2.001b 0.2 0.2 1.0 0.67 1.493 0.2 0.2 0.2 0.25 0.803a 0.2 0.2 0.2 0.50 0.403b 0.2 0.2 0.2 0.67 0.304 0.5 0.5 2.5 0.25 10.04a 0.5 0.5 2.5 0.33 7.584b 0.5 0.5 2.5 0.40 6.256 0.5 0.5 0.5 0.25 2.006a 0.5 0.5 0.5 0.33 1.526b 0.5 0.5 0.5 0.40 1.25
-
40
3.5 Mesh Refinement
Two sets of finite element models were constructed using the McClung et
al. geometries [15] found in Table 3.3. The first set contained a coarse mesh refinement
along the crack front. The coarse mesh refinement along the crack front can be seen
in Figure 3.5. The second set of models had three times more elements around the
crack front (Figure 3.14). Table 3.5 shows the number of crack front nodes in the
coarse and refined meshes.
3.6 Finite Size Effects
FEMs were generated to test the effect of specimen height and width on the
J-integral. The a/t ratios of 0.2 and 0.5, and the a/c ratios of 0.2 and 1.0 were
used in this analysis. The height effect models utilized the crack ratios for Model 1
(a/t = 0.2, a/c = 0.2), Model 4 (a/t = 0.5, a/c = 0.2), and Model 9 (a/t = 0.8, a/c =
1.0). The width effect models utilized the same model geometries used in the Nasgro
J-comparison work (Table 3.4).
Table 3.5 Number of crack front nodes in the coarse and refined meshesModel a/t a/c Coarse Refined
1 0.2 0.2 31 912 0.2 0.6 17 493 0.2 1.0 17 494 0.5 0.2 45 1335 0.5 0.6 17 496 0.5 1.0 17 497 0.8 0.2 73 2658 0.8 0.6 31 919 0.8 1.0 17 49
-
41
Figure 3.14 Refined mesh along the crack front
3.7 Material Properties
The material properties, unless otherwise specified, were based on a structural
steel. These are the same material properties used Natarajan [22] for some FEMs
in his thesis work involving J-integral solutions. Two different yielding models were
used in this research. The first was the Ramberg-Osgood deformation plasticity
model. The second was an incremental plasticity method requiring a table of σ and
εpl. The elastic material properties for each model depended on the yielding scheme
used for the FEA.
3.7.1 Deformation Plasticity
The following material properties were used with the *Deformation Plasticity
command in ABAQUS:
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42
• E = 30.0× 106 psi
• ν = 0.3
• σo = 40.0× 103 psi
• α = 0.5
• n = 5, 10, and 15
where E is Young’s modulus, ν is Poisson’s ratio, σo is yield or reference stress, α
is a dimensionless constant as described in Equation 2.4, and n is the hardening
exponent. The effect of n on the stress vs. strain curves modelled using the Ramberg-
Osgood equation is shown in Figure 3.15. Notice that the smaller n is, the greater
the hardening slope
Figure 3.15 Effect of n on the stress vs. strain curve using a Ramberg-Osgoodmodel
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43
3.7.2 Incremental Plasticity
The incremental plasticity models, with the exception of the Kirk and Dodds
comparison work, were generated using the Ramberg-Osgood equation,
ε
εo=
σ
σ0+ α
(σ
σ0
)n, (3.10)
shown again for convenience. The material properties listed in the previous section
were used to generate the a new Young’s modulus and a table of stress vs. plastic
strain for use in ABAQUS. The Young’s modulus, E = 30.0 × 106 psi, used for the
fully plastic analyses was not used to derive the stress vs. plastic strain tables for
ABAQUS. It was replaced by a secant modulus,∗E, as shown in Equation 3.11:
∗E =
σ0εo (1 + α)
. (3.11)
This value of Young’s modulus was selected because it intersects the Ramberg-Osgood
curve at the fully plastic reference stress, σo = 40.0× 103 psi (Figure 3.16).
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44
Figure 3.16 Intersection of Ramberg-Osgood curves at σo
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45
Using this scheme, the elastic strain, and therefore the J-integral, will be
underestimated at low stresses (Figure 3.17). But, for sufficiently high stresses, the
elastic strain becomes overwhelmed by the plastic strain, making the error negligible.
The reference strain can now be expressed as
εo =σ0∗E
. (3.12)
Figure 3.17 Elastic, modified elastic, and Ramberg-Osgood stress vs. strain curvesfor n = 10
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46
Multiplying both sides of Equation 3.10 by εo and substituting Equation 3.12 yields:
ε =σ∗E
+ αεo
(σ
σ0
)n. (3.13)
Equation 3.13 can be divided into the elastic and plastic strains as
εel =σ∗E
, (3.14)
and
εpl = αεo
(σ
σo
)n. (3.15)
The plastic strains at different stresses were then calculated for use with the *PLAS-
TIC command in ABAQUS for incremental plasticity analyses.
In summary, elastic-plastic material properties used in this research are based
on a modified Young’s modulus. This modification makes it possible to generate
incremental plasticity models that exhibit the same yield stress for all n’s. The elastic
properties used for the incremental plasticity analyses are∗E = 20× 106 and ν = 0.3.
The stress vs. plastic strain values used with the *Plastic command in ABAQUS are
shown in Tables 3.6 - 3.8.
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Table 3.6 Stress vs. plastic strain data at n = 15, used for ABAQUS models
Stress Plastic Strain40000 0.00066741200 0.00103942400 0.00159843600 0.00242844800 0.00364946000 0.00542547200 0.00798248400 0.01163349600 0.01679750800 0.02404252000 0.03412453200 0.04804954400 0.06714255600 0.09313956800 0.12830258000 0.17556159200 0.2386960400 0.32252561600 0.43323162800 0.57863364000 0.76861465200 1.0156
Table 3.7 Stress vs. plastic strain data at n = 10, used for ABAQUS models
Stress Plastic Strain40000 0.044000 0.00172916248400 0.00448552800 0.01070651357200 0.02383796261600 0.05001680566000 0.09971217470400 0.19012333374800 0.34859793279200 0.61739149783600 1.060160459
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Table 3.8 Stress vs. plastic strain data at n = 5, used for ABAQUS models
Stress Plastic Strain40000 0.044000 0.00107448400 0.00172952800 0.00267257200 0.00398661600 0.00577466000 0.00815370400 0.01125874800 0.01524579200 0.02028883600 0.02658588000 0.03435892400 0.0438596800 0.055333101200 0.069105105600 0.085493110000 0.104851114400 0.127567118800 0.15406123200 0.184783127600 0.220223132000 0.260903136400 0.307384140800 0.360265145200 0.420186149600 0.487828154000 0.563913158400 0.649209162800 0.744528167200 0.850727171600 0.968713176000 1.099441
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CHAPTER 4
RESULTS
This chapter begins with a discussion of the results for various fully plastic
element sets. Next, models generated for parameters used by Kirk and Dodds [23]
are compared to published results. McClung, Lei, and Newman-Raju data are then
compared to current FEM results. Finally, the effects of the specimen size on the
J-integral are examined, and the h1 values for various specimen widths are compared
with Nasgro results.
4.1 Fully Plastic Zone
The mesh generation program mesh3d scp was used to generate FEM’s for all
four of the fully plastic zones described in Chapter 3. The same mesh was used for
each model. Only the specified fully plastic element set was changed for the different
models. It was found that the J-integral was identical for all of the described zones.
Therefore, only Partial Layer 1 was used in later fully plastic models was used for
the FEA-Crack meshes, and LayerCR was used for any fully plastic meshes produced
with mesh scp.
4.2 Kirk and Dodds Incremental Plasticity
The results for models generated per the Kirk and Dodds geometries are shown
in Table 4.1. The results compared quite well to the published data. The maximum
difference between the current results and the published data was 2.9%. It should be
noted that this excellent agreement in results was obtained even though the meshes
49
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50
Table 4.1 Comparison of FEM results to Kirk and Dodds values
Crack θ◦ J (in-lb) Kirk and Dodds % Differencefrom FEM J (in-lb)
1 30.9 0.749 0.732 2.31 90 0.892 0.867 2.92 30.9 2.055 2.014 2.02 90 0.892 0.867 2.933 30.9 2.077 2.046 1.53 90 3.207 3.173 1.7
used by Kirk and Dodds contained approximately 25% the number of nodes and
elements used in this research.
4.3 McClung and Lei Comparisons
Elastic, fully plastic, and incremental plasticity FEA results for the McClung
et al. geometries are presented in this section. The elastic results are compared to the
Newman-Raju [24] calculations, and graphical trends are noted in the comparison of
Lei’s [17] elastic results. The fully plastic data are compared to the tabular data of
McClung et al. [15] and Lei [17]. The effects of mesh refinement are discussed for
both the elastic and fully plastic FEMs. Finally the incremental plasticity and fully
plastic FEA results are compared.
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51
Figure 4.1 Model 1 (a/t=0.2, a/c=0.2): Normalized K factor vs. angle along crackfront
4.3.1 Elastic Analysis
The K factors obtained from the FEM’s with the McClung geometries were
normalized using
Knorm =KI
σ√
π aQ
, (4.1)
from Newman and Raju [25]. The results of the elastic FEM models and the Newman
and Raju [24] calculations are presented in Figures 4.1-4.9.
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52
Figure 4.2 Model 2 (a/t=0.2, a/c=0.6): Normalized K factor vs. angle along crackfront
Figure 4.3 Model 3 (a/t=0.2, a/c=1.0): Normalized K factor vs. angle along crackfront
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53
Figure 4.4 Model 4 (a/t=0.5, a/c=0.2): Normalized K factor vs. angle along crackfront
Figure 4.5 Model 5 (a/t=0.5, a/c=0.6): Normalized K factor vs. angle along crackfront
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54
Figure 4.6 Model 6 (a/t=0.5, a/c=1.0): Normalized K factor vs. angle along crackfront
Figure 4.7 Model 7 (a/t=0.8, a/c=0.2): Normalized K factor vs. angle along crackfront
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55
Figure 4.8 Model 8 (a/t=0.8, a/c=0.6): Normalized K factor vs. angle along crackfront
Figure 4.9 Model 9 (a/t=0.8, a/c=1.0): Normalized K factor vs. angle along crackfront
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56
Significant increases in the K-factor at the surface and/or depth were observed
in all of the models. Only Models 8 (a/t = 0.8, a/c = 0.6) and 9 (a/t = 0.8, a/c = 1.0)
do not have large increases in the K-factor at the free surface. Models 1, 4, and 7
(all with a/c = 0.2) are the only FEMs that do not contain the same K-factor spike
repeated in the depth (Table 4.2).
When the surface and depth spikes are disregarded, the ABAQUS results com-
pared very reasonably to the normalized K-factors calculated per the Newman and
Raju [24] equations. This favorable comparison occurred even though the mid-side
nodes were not moved to the quarter points, and the nodes along the crack tip were
left untied (two conditions which yield optimum accuracy in K-factor calculations
using FEMs). The largest observed error, approximately six percent, occurred with
Model 8. It should also be noted that the normalized K-factor results from the K
and elastic J models were very close. The elastic results are summarized in Table
4.3. This summary disregards the surface and depth results. There is no apparent
pattern to the differences.
The current FEA results were also compared visually to the graphical results
published by Lei [17]. Lei used a different normalizing scheme, resulting in different
Table 4.2 Surface and depth phenomenon for K-factors
Surface DepthModel a/t a/c Jump Jump
1 0.2 0.2 yes no2 0.2 0.6 yes yes3 0.2 1.0 yes yes4 0.5 0.2 yes no5 0.5 0.6 yes yes6 0.5 1.0 yes yes7 0.8 0.2 yes no8 0.8 0.6 no yes9 0.8 1.0 no yes
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57
Table 4.3 Maximum percent differences between Newman-Raju and FEM solu-tions (quarter symmetry)
Max FEA K Max FEA K Max FEA KDirectly from Jel from tied nodes
Model a/t a/c (% diff.) (% diff.) (% diff.)1 0.2 0.2 -5.64 -5.38 -8.52 0.2 0.6 -2.28 -2.18 -3 0.2 1.0 3.11 2.88 -4 0.5 0.2 -6.40 -5.57 -5 0.5 0.6 -2.15 -1.96 -6 0.5 1.0 3.35 3.44 -7 0.8 0.2 4.52 4.53 -8 0.8 0.6 -7.07 -7.07 -7.179 0.8 1.0 2.8 3.03 -
scales on the y-axis, but the graphs had very similar shapes. Lei also showed some
models with the same spike at the surface that was experienced in this research.
However, the increase was not as significant. No sudden increases were observed at
the depth of his elastic models.
An investigation was performed to find the cause of the previously discussed
surface and depth K-factor spikes in the current FEMs. The first step was to explore
the potential error caused by not tying the crack tip nodes or moving the mid-side
nodes to the quarter points. Model 1 (a/t = 0.2, a/c = 0.2) and Model 8 (a/t = 0.8,
a/c = 0.6) meshes were recreated in FEA-Crack for elastic analysis only. These two
models were generated with an elastic singularity, 1/√
r, created by tying the crack
tip nodes and moving the mid-side nodes to the quarter points (Figure 4.10). The
results of the two elastic models are shown graphically in Figures 4.11 and 4.12. The
K-factors produced using these two FEM’s were almost identical to the previous
results for Model 1 and Model 8.
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58
Figure 4.10 Elastic singularity element [4]
Figure 4.11 Model 1 (a/t=0.2, a/c=0.2): Normalized K-factor vs. angle alongcrack front for untied and tied nodes
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59
Figure 4.12 Model 8 (a/t=0.8, a/c=0.6): Normalized K-factor vs. angle alongcrack front for untied and tied nodes
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60
The second step in investigating the surface and depth K-factor spikes in-
volved changing from reduced integration to full integration elements, type C3D20R
to C3D20 in ABAQUS, for the FEMs. The results for Models 1 and 8 are shown in
Figures 4.13 and 4.14. As mentioned in the ABAQUS User’s Manuals [21], using a
full integration element type caused the K-factor results to oscillate. Full integration
elements did not correct the surface and depth deviations.
Figure 4.13 Model 1 (a/t = 0.2, a/c = 0.2): Reduced vs. full integration elements
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61
Figure 4.14 Model 8 (a/t = 0.8, a/c = 0.6): Reduced vs. full integration elements
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62
The default nonlinear solver for ABAQUS was also considered as a possible
source of error in the third attempt to reduce the surface and depth K-factor dis-
crepencies. Since the K-factor calculation is linear, the solver used within ABAQUS
was changed to a linear perturbation. Unfortunately, the *CONTOUR INTEGRAL
command used to output the K-factors will not function within a linear perturbation
step. Another attempt to force a linear solution was made by using the *STATIC
command to force the solution to be performed in one step. The surface and depth