Citation
Liquefaction resistance of Monterey no.0/30 sand containing fines under cyclic triaxial and cyclic hollow cylinder tests

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Title:
Liquefaction resistance of Monterey no.0/30 sand containing fines under cyclic triaxial and cyclic hollow cylinder tests
Creator:
Liu, Jungang
Place of Publication:
Denver, CO
Publisher:
University of Colorado Denver
Publication Date:
Language:
English

Thesis/Dissertation Information

Degree:
Doctorate ( Doctor of philosophy)
Degree Grantor:
University of Colorado Denver
Degree Divisions:
Department of Civil Engineering, CU Denver
Degree Disciplines:
Civil engineering
Committee Chair:
Dashti, Shideh
Committee Members:
Chang, Nien-Yun
Znidarcic, Dobroslav
Brady, Brian
Hien, Nghiem
Wang, Shingchun

Notes

Abstract:
Liquefaction is the most detrimental ground failure caused by strong earthquakes. Ground liquefaction leads to associated foundation and superstructure failures due to loss of bearing capacity and excessive deformation and an appropriate assessment of liquefaction is critical to the seismic safety evaluation of foundations and super structures. This study concerns the liquefaction resistance of Monterey No. 0/30 sand contacting fines. Monterey No. 0/30 sand is a uniform clean medium sand. Fines were prepared by sieving Leyden clay through a U.S.# 200 sieve. One hundred fourteen isotopically consolidated undrain cyclic triaxial tests and thirtyseven cyclic hollow cylinder test were performed to investigate the effect of fines content on liquefaction resistance of soils. This research includes compare and relate the soil liquefaction resistance found by cyclic triaixal and cyclic hollow cylinder test results on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content. The cyclic triaxial and cyclic hollow cylinder liquefaction resistances of soils were expressed in terms of liquefaction potential curves. From the liquefaction potential curves, stress ratio in cyclic triaxial and cyclic hollow cylinder tests causing initial liquefaction in 10 cycles, 30 cycles, 40 cycles and 50 cycles were chosen as dependent variables in the regression model. Three independent variables of regression models, deviator stress in cyclic triaxial test, DS (cyclic shear stress in cyclic hollow cylinder test, CSS); fine content (decimal), FC; consolidation pressure, CP, were eventually selected for final statistical analysis. In addition to the evaluation of liquefaction potential, an excess pore pressure generation was simulated using Horita constitutive model with the parameters evaluated by laboratory tests. This research includes comparison of excess pore pressure generation between simulations from Horita’s constitutive model and measured from cyclic triaxial test. The new findings of this study can be applied to better understand the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils.

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University of Colorado Denver
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Auraria Library
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Copyright Jungang Liu. Permission granted to University of Colorado Denver to digitize and display this item for non-profit research and educational purposes. Any reuse of this item in excess of fair use or other copyright exemptions requires permission of the copyright holder.

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Full Text
LIQUEFACTION RESISTANCE OF MONTEREY NO.0/30 SAND CONTAINING FINES UNDER CYCLIC TRIAXIAL AND CYCLIC HOLLOW CYLINDER TESTS
by
JUNGANG LIU
B.S., Wuhan University of Technology, China, 2007 M.S., University of Colorado Denver, 2012
A thesis submitted to the Faculty of the Graduate School of the University of Colorado Denver in partial fulfillment of the requirements for the degree of Doctor of Philosophy Civil Engineering Program 2019


©2019
JUNGANG LIU ALL RIGHTS RESERVED


This thesis for the Doctor of Philosophy degree by Jungang Liu
Has been approved for the Civil Engineering Program by
Shideh Dashti, Chair Nien-Yin Chang, Advisor Dobroslav Znidarcic Brian Brady Nghiem Hien Shingchun Wang
Date: May, 18 2019


Liu, Jungang (Ph. D, Civil Engineering)
Liquefaction Resistance of Monterey No. 0/30 Sand Containing Fines under Cyclic Triaxial and Cyclic Hollow Cylinder Tests
Thesis directed by Professor Nien-Yin Chang
ABSTRACT
Liquefaction is the most detrimental ground failure caused by strong earthquakes. Ground liquefaction leads to associated foundation and superstructure failures due to loss of bearing capacity and excessive deformation and an appropriate assessment of liquefaction is critical to the seismic safety evaluation of foundations and super structures. This study concerns the liquefaction resistance of Monterey No. 0/30 sand contacting fines. Monterey No. 0/30 sand is a uniform clean medium sand. Fines were prepared by sieving Leyden clay through a U.S.# 200 sieve. One hundred fourteen isotopically consolidated undrain cyclic triaxial tests and thirty-seven cyclic hollow cylinder test were performed to investigate the effect of fines content on liquefaction resistance of soils. This research includes compare and relate the soil liquefaction resistance found by cyclic triaixal and cyclic hollow cylinder test results on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content.
The cyclic triaxial and cyclic hollow cylinder liquefaction resistances of soils were expressed in terms of liquefaction potential curves. From the liquefaction potential curves, stress ratio in cyclic triaxial and cyclic hollow cylinder tests causing initial liquefaction in 10 cycles, 30 cycles, 40 cycles and 50 cycles were chosen as dependent variables in the regression model. Three independent variables of regression models, deviator stress in cyclic triaxial test, DS (cyclic shear stress in cyclic hollow cylinder test, CSS); fine content (decimal), FC; consolidation pressure, CP, were eventually selected for final statistical analysis.


In addition to the evaluation of liquefaction potential, an excess pore pressure generation was simulated using Horita constitutive model with the parameters evaluated by laboratory tests. This research includes comparison of excess pore pressure generation between simulations from Horita’s constitutive model and measured from cyclic triaxial test. The new findings of this study can be applied to better understand the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils.
The form and content of this abstract are approved. I recommend its publication.
Approved: Nien-Yin Chang


ACKNOWLEDGEMENTS
I would like to express my sincerest appreciation to my advisor, Professor Nien-Yin Chang for his constant support and guidance throughout the course of this study. Gratitude is extended to Dr. Shideh Dashti, Dr. Dobroslav Znidarcic, Dr. Brian Brady, Dr. Nghiem Hien and Dr. Shingchun Wang for serving on my final examination committee.
I would like to thank my student colleague, Mr. Brian Volmer for his friendship and help in preparing this study. Thanks are extended to Mr. Tom Thuis and his staff in Calibration shop at the University of Colorado Denver for their assistance in instrumentation.
I would like to thank my wife, Liang Feng for many years of her endurance, sacrifice, assistance and encouragement, which made this accomplishment possible. I also thank my parents for their understanding and support.


TABLE OF CONTENTS
CHAPTER
1. INTRODUCTION................................................................1
Problem Statement...........................................................1
Objectives and Scope........................................................2
Significance of This Research...............................................3
2. LITERATURE REVIEW: LIQUEFACTION RESISTANCE..................................4
Liquefaction Resistance from Earthquake Case History Data...................4
1964 Niigata Earthquake...............................................4
1964 Alaska Earthquake................................................6
1995 Hyogoken-Nanbu Earthquake........................................7
1999 Chi-Chi Earthquake..............................................10
2011 Tohoku Earthquake...............................................11
2011 Christchurch Earthquake.........................................13
Field Methods for Soil Liquefaction Resistance Evaluation..................18
General..............................................................18
Standard Penetration Test (SPT)......................................19
Cone Penetration Test (CPT)..........................................31
Piezometric Cone Penetrometer Test...................................41
Seismic Cone Penetration Test........................................41
Other Techniques (Appendix A)........................................44
Laboratory Methods for Soil Liquefaction Resistance Evaluation.............45
Undisturbed Sampling
45


Laboratory Test
46
Cyclic Triaxial Tests...............................................46
Cyclic Hollow Cylinder Tests........................................50
Cyclic Simple Shear Test............................................55
Shaking Table Test..................................................57
Centrifuge Test....................................................57
Factor Safety against Liquefaction................................................58
3. LITERATURE REVIEW: FACTORS AFFECTING LIQUEFACTION RESISTANCE
OF SOILS.........................................................................61
General..........................................................................61
Effects of Soil Factors..........................................................61
Soil Properties...........................................................61
Fines Contents and Plasticity Index.......................................64
Soil Relative Density.....................................................72
Particle Size Gradation...................................................73
Particle Shape............................................................73
Geological Aging and Cementation..........................................73
Effects of Laboratory Factors....................................................75
Specimen Preparation Method...............................................75
Reconstitution versus Intact Specimens....................................76
Loading Wave Forms........................................................76
Frequency on Cyclic Strength..............................................78
Specimen Size.............................................................78


Frictionless Caps and Bases..............................................79
Membrane Compliance......................................................79
Relative Density.........................................................80
Confining Stress (03)....................................................81
Cyclic Stress Amplitude and Number of Cyclic Stress Cycles...............82
Particles Size and Gradation.............................................82
Pre-straining............................................................83
Lateral Earth Pressure (Ko) and Over Consolidation Ratio.................83
Consolidation Ratio, (Kc)................................................84
Effects of Field Factors........................................................85
Groundwater Table........................................................85
Placement Conditions or Depositional Environment.........................86
Drainage Condition.......................................................86
Confining Pressure.......................................................86
Historical Environment...................................................87
Building Load............................................................88
Effects of Earthquake Factors...................................................88
Magnitude Scaling........................................................88
Intensity and Duration...................................................91
Ground Motions...........................................................92
4. CYCLIC TRIAXIAL TEST PROGRAM AND RESULTS ANALYSIS...............................98
Introduction....................................................................98
Test Program....................................................................99


Test Equipment................................................................102
Soil Samples Preparation......................................................105
Soil Samples...........................................................105
Mixing Soil Samples....................................................105
Samples Preparation....................................................107
Samples Quality Control................................................123
Relative Density Control.........................................123
Soil Samples Saturation..........................................123
Determining B-Parameter..........................................124
Test Procedure................................................................124
Tests Results.................................................................128
Soil Liquefaction Potential Curves.....................................129
Cyclic Axial load versus Number of Cycles to liquefaction..............133
Excess Pore Water Pressure versus Number of Cycles to liquefaction.....134
Cyclic Deviator Stress versus Axial Strain.............................135
Stress Path............................................................136
Effect of Fines Content on Liquefaction Resistance.....................137
5. HOLLOW CYLINDER TEST APPARATUS, PROGRAMS AND TEST RESULTS
ANALYSIS......................................................................138
Background....................................................................138
Hollow Cylinder Test Apparatus................................................139
Principles of Hollow Cylinder Testing..................................139
Stress Distribution in Hollow Cylinder Specimens.......................143


Specimen Geometry......................................................147
Membrane Penetration Errors............................................149
Soil Samples Preparation......................................................153
Soil Samples...........................................................153
Mixing Soil Samples....................................................153
Samples Preparation....................................................153
Samples Quality Control................................................158
Relative Density Control.........................................158
Soil Samples Saturation..........................................159
Determining B-Parameter..........................................159
Tests Program.................................................................159
Test Procedure................................................................174
Analysis of Results...........................................................176
Cyclic Torque versus Number of Cycles to liquefaction..................177
Excess Pore Water Pressure versus Number of Cycles to liquefaction.....178
Cyclic Shear Stress versus Shear Strain................................179
Stress Path............................................................180
Effect of Fines Content on Liquefaction Resistance.....................181
6. COMPARISON OF CYCLIC TRIAXIAL AND HOLLOW CYLINDER TEST
RESULTS.......................................................................182
Introduction..................................................................182
Cyclic Strength...............................................................185
Correction Factor between Cyclic Stress Ratio Causing Liquefaction in the


Field and Cyclic Stress Ratio Causing Liquefaction of Triaxial Test Sample in the
Laboratory..............................................................185
Comparison on both test results................................................190
Cyclic Torque, Cyclic Axial Load versus Number of Cycles to Liquefaction... 190
Pore Water Pressure versus Number of Cycles to Liquefaction.............193
Cyclic Shear Stress, Cyclic Deviator Stress versus Shear Strain, Axial Strain... 196 Stress Path.............................................................199
7. THRESHOLD FINE CONTENTS........................................................202
Previous Studies...............................................................202
Threshold Fines Content........................................................212
Definition..............................................................212
Factors Effects Threshold Fines Content.................................212
Results of laboratory Testing..................................................217
All Soil Samples Results in Cyclic Triaxial Test........................217
All Soil Samples Results in Cyclic Hollow Cylinder Test.................220
8. EXCESS PORE PRESSURE GENERATION................................................225
Excess pore pressure generation from laboratory tests..........................225
Previous Studies........................................................225
Excess pore pressure Generation from laboratory test results............230
Constitutive models for simulating pore pressure generation....................236
Constitute Model UBC3D-PLM..............................................236
Finn Constitutive model.................................................238
Horita’s model.................................................................239


9. STATISTICAL MODELING OF LIQUEFACTION RESISTANCE
250
Introduction.............................................................250
Selection of Variables...................................................250
Initial Variable Selection........................................250
Variable Reduction................................................251
Final Variable Selection..........................................254
Regression Model for Liquefaction Resistance.............................263
Regression model for cyclic triaxial test.........................263
Regression model for cyclic hollow cylinder test..................265
10. A PROPOSED NEW PROCEDURE FOR EVALUATING LIQUEFACTION
RESISTANCE OF SOIL WITH PLASTIC FINES....................................266
Introduction.............................................................266
Procedure by Seed et al..................................................266
Proposed Procedure Cyclic Resistance Ratio from SPT based case history...270
SPT-based Case History Database from Idriss and Boulanger.........271
Cases History from Kohji Tokimatsu and Yoshiaki Yoshimi...........278
Proposed Procedure Cyclic Resistance Ratio from SPT based laboratory testing data...286
Calculating SPT Blow Count (Ni)60 Procedure.......................286
Calculation of SPT Blow Count based Laboratory Test Data..........289
Comment on the New Procedure.............................................308
11. SUMMARY CONCLUSION AND RECOMMENDATIONS FOR FUTURE
STUDIES
310


Summary
310
Conclusions.............................................................311
Recommendations for Future Studies......................................312
REFERENCES.....................................................................313
APPENDIX.......................................................................325
A. Field Methods for Soil Liquefaction Resistance Evaluation...............325
B. Factors Affecting Liquefaction Resistance of Soils......................335
C. Procedure for Fabricating Rubber Membranes..............................346
D. Cyclic Triaxial and Hollow Cylinder Tests Results.......................349


LIST OF TABLES
TABLE
1. Summary of CPT Field case histories from the 1995 Hyogoken-Nambu (Kobe) earthquake.
(Moss, 2003) ............................................................................9
2. Major earthquake case histories (Many researchers)......................................16
3. Magnitude correction factors for cyclic stress approach (Seed et al, 1975).............25
4. Relationship between (Ni)6o and potential damage. (Seed et al, 1985)..................25
5. Relationship between fines content and tip resistance increment (Seed et al, 2003).....38
6. Estimated susceptibility deposits to liquefaction during strong seismic shaking based on
geologic age and depositional environment. (Many researchers)..........................96
7. All soil samples in cyclic triaxial test (Jungang Liu, 2019)...........................128
8. Test program details in hollow cylinder test (Jungang Liu, 2019).......................176
9. Result of comparative study on clean sand (Liu, et al, 2017)...........................183
10. Result of comparative study on mixture sample (Jungang Liu, 2019).....................184
11. Values of Cr (some researchers).......................................................185
12. Cr values on the clean sand produced in Laboratory by CTT and CHCT (Liu, et al,
2017).................................................................................187
13. Cr values on the mixture samples produced in laboratory by CTT and CHCT (Jungang Liu,
2019).................................................................................188
14. Correlation coefficient matrix of dependent variable Cr’ and independent variables....189
15. Threshold fines content in cyclic triaxial test and cyclic hollow cylinder test (Jungang Liu,
2019).................................................................................217


16. Correlation coefficient matrix of threshold fines content and laboratory factors (Jungang Liu,
2019)................................................................................217
17. All parameters in Horita model for trail 1 (Masakuni, Horita, 1985).................247
18. All parameters in Horita model for trail 2 (Masakuni, Horita, 1985).................248
19. Correlation coefficient matrix of the dependent variable, SRCTT10 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)...................................255
20. Correlation coefficient matrix of the dependent variable, SRCTT30 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)...................................256
21. Correlation coefficient matrix of the dependent variable, SRCTT40 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)...................................257
22. Correlation coefficient matrix of the dependent variable, SRCTT50 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)....................................258
23. Correlation coefficient matrix of the dependent variable, SRCTT30 (Dr30) and the sixteen
chosen primary independent variables (Jungang Liu, 2019)............................259
24. Correlation coefficient matrix of the dependent variable, SRCTT30 (ems) and the sixteen
chosen primary independent variables (Jungang Liu, 2019).............................260
25. Correlation coefficient matrix of the dependent variable, SRCHCT10 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)....................................261
26. Correlation coefficient matrix of the dependent variable, SRCHCT30 and the sixteen chosen
primary independent variables (Jungang Liu, 2019)....................................262
27. Summary of SPT-based liquefaction case history data from Idriss and Boulanger (2004,
2008)................................................................................274


28. Case histories data from Idriss and Boulanger: (Ni)6o and CSR with different Fines Content
data were created by Jungang Liu (2019)..............................................277
29. Summary of SPT-based liquefaction case history data from Tokimatsu and Yoshimi
(1983)...............................................................................283
30. Case histories data from Idriss and Boulanger: Ni and CSR with different Fines Content data
were created by Jungang Liu (2019)...................................................285
31. Summary of Rod Energy Ratios (Skempton, A.W. 1986)..................................288


LIST OF FIGURES
FIGURE
2.1 Niigata city destroyed by the Niigata Earthquake 1964. (National Geophysical Data Center).5
2.2 Soil profile at Kawagishi-cho (Niigata Prefecture) (National Geophysical Data Center).6
2.3 Cracked Highway in 1964 Alaska Earthquake (Timothy J. Walsh, et. al 1995).............7
2.4 Road Embankment Failure Caused by Soil Liquefaction. (Timothy J. Walsh, et. al 1995)..7
2.5 Location of aftershocks (Moss, 2003)..................................................8
2.6 Map Showing CPT Field Case History Locations from the 1995 Hyogoken-Nambu
Earthquake, in relation to the SPT Case History Locations and PGA Contours from Cetin et al. (2000)......................................................................9
2.7 Proposed CPT-based liquefaction boundary curve (Ku, Chih-Sheng, Der-Her Lee, and Jian-
Hong Wu 2003)........................................................................11
2.8 Observed liquefaction in the park (left) and parking area of Disneyland (right). (S.
Bhattacharya, et, 2011)................................................................12
2.9 Typical soil profile in Urayasu (S. Bhattacharya, et, 2011)...........................13
2.10. Ground Failures and Resulting Water Damage. (Christchurch, New Zealand Earthquake Field Report, 2011)................................................................15
2.11 Liquefaction documentation map of eastern Christchurch from drive-through
Reconnaissance. (Christchurch, New Zealand Earthquake Field Report, 2011)...........15
2.12 Relationship between cyclic stress ratios causing liquefaction and (Ni)6o values for clean
Sands inM= 7.5 earthquakes. (After Seed et al. (1975).................................23
2.13 Relationship between cyclic stress ratios causing liquefaction and (Ni)6o values for silty
Sands inM= 7.5 earthquakes. (After Seed et al. (1975).................................24
2.14 Updated SPT case history database of liquefaction in cohesionless soils with various fines
Contents in terms of equivalent CSR for M = 7.5 and o'v = 1 atm and equivalent clean sand (Ni)60cs (Idriss-Boulanger, 2008).............................................27
2.15 SPT case history database used previously by Idriss and Boulanger (2004, 2008).......28
2.16 Parts (a) and (b) of Figure 2.6 in Cetin et al. (2004); note that the points representing case
Histories are identical in parts (a) and (b) of this figure and as representing conditions with o'v= 1 atm....................................................................29


30
2.17 Curves relating CSR to (Ni)6o published over the past 24 years for clean sands and the
Recommended curve for M = TA and o'V0= 1 atm (Seed et al 2001)..............
2.18 Manufacturing and operating tolerances of cones, taken from ASTM D5778.......33
2.19 Schematic section through a piezocone head, showing the piezo-element and friction sleeve
(From ASTMD5778)...............................................................34
2.20 CPT prediction of overburden pressure-corrected SPT blow-count (Olsen and Malone,
2000)..........................................................................35
2.21 CPT-based case histories and recommended relation for clean sands for M =7*/2 and 0V0 = 1
atm (I.M. Idriss and R.W. Boulanger, 2004).....................................39
2.22 CPT-based case histories and recommended relation for clean sands with relations
Proposed by others.............................................................40
2.23 Seismic Cone ( from A.P. van den berg)........................................43
2.24 Seismic cone configuration (from A.P. van den berg)...........................43
2.25 Results of the SCPT carried out on June 16th 2001 in a permafrost mound near Umiujaq,
Nunavik, Canada................................................................44
2.26 Schematic of cyclic triaxial test equipment (Marcuson and Krinitzsky, 1976)...48
2.27 Typical analog recordings of load, deformation, and pore pressures during a cyclic triaxial
Test (Department of the Army, 1990)............................................49
2.28 Cyclic triaxial strength curves for Monterey No.O sand (Department of the Army, 1990). ..50
2.29 Idealized stress and strain components within the HCA subjected to axial load, W, torque, Mr, internal pressure, Pi, and external pressure, Po: (a) hollow cylinder coordinates; (b) Element component stresses; (c) element component strains; (d) element principal stresses
(After Zdravkovic and Jardine, 2001).................................................52
2.30 Cyclic Simple Shear Test Device.....................................................56
2.31 Reduction factor to estimate the variation of cyclic shear stress with depth below level or
Gently sloping ground surfaces. (After Seed and Idriss, 1971)........................60
3.1 Data presented by Wang (1979) which led to the development of the Chinese Criteria....63
3.2 Characteristics of Adapazari soils that were tested in J.D.Bray & R.B.Sancio (2006)’s study:
(a) Location on the Casagrande plasticity chart; (b) relationship between liquid limit and Moisture content...................................................................63


3.3 SPT case histories of cohesionless soils with FC > 35% and the NCEER Workshop (1997) Curve and the recommended curves for both clean sand and for FC = 35% for M = I'A and o
Vo = 1 atm (FM.Idriss and R.W.Boulanger, 2004).......................................66
3.4 SPT case histories of cohesionless soils with 5% Both clean sand and for FC = 15% for M = I'A and o Vo = 1 atm (FM.Idriss and R.W.Boulanger, 2004)..............................................................67
3.5 SPT case histories of cohesionless soils with 15% (1997) curve and the recommended curves for both clean sand and for FC = 15% for M = I'A And o Vo= 1 atm (FM.Idriss and R.W.Boulanger, 2004)...............................68
3.6 Stress Ratio required to reach Liquefaction in 10 Cycles versus Fine Content Under
Confining pressure 15 psi (N.Y. Chang, 1990).........................................69
3.7 Stress Ratio Required to reach Liquefaction in 30 Cycles versus Fine Content
Under Confining pressure 15 psi (N.Y.Chang, 1990)....................................69
3.8 Stress Ratio Required to reach Liquefaction in 100 Cycles versus Fine Content Under
Confining pressure 15 psi (N.Y.Chang, 1990)..........................................70
3.9 Cyclic resistance ratio versus equivalent (FC)400 (Tzuo-Shin Ueng, Chia-Wen Sun and
Chieh-When Chen, 2013)................................................................70
3.10 Variation in Liquefaction Resistance with Fines Content for Ni=20 (Yong Wang and Yanli
Wang, 2010)..........................................................................71
3.11 Change of relative density and liquefaction potential with different fines contents and Silts
For tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011).............71
3.12 Change of void ratio and liquefaction potential with different fines contents and silts for
Tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011).................72
3.13 Effect of loading wave form on cycles to initial liquefaction for moist-tamped Specimen.
(Silver, et al, 2000)................................................................78
3.14 Magnitude scaling factor (MSF) relationship by some researchers.....................89
3.15 MSF relationships for clay and sand (Boulanger and Idriss 2007).....................90
3.16 Variation in the MSF relationship with qciNcs and with (Nl)60cs for cohesionless soils
(Boulanger and Idriss 2007)..........................................................91
3.17 Approximate relationship between maximum accelerations on rock and maximum ground
Accelerations (Seed et al., 1976)....................................................94


3.18 Amplification - attenuation relationship for modifying bedrock acceleration at soft soil sites (Idriss, 1990).........................................................................95
4.1 Test Specimen Placed Within the Triaxial Cell. (Jungang Liu, 2019)................100
4.2 Pore pressure transducer and a 4-way valve used to monitor cell pressure, back pressure, and
Specimen pore water pressure. (Jungang Liu, 2019).....................................101
4.3 Pressure control panel used to apply cell and back pressure to the triaxial cell. Graduated
Burettes were used to determine specimen volume change. (Jungang Liu, 2019)...........101
4.4 Closed loop electro-hydraulic materials test system applying a sinusoidal loading to
Soil specimen. (Jungang Liu, 2019)................................................102
4.5 Rubber o-ring (Jungang Liu, 2019)................................................115
4.6 Vernier caliper used to measure the thickness of the membrane and diameter of the sand
Specimen (Jungang Liu, 2019)......................................................115
4.7 Vernier caliper used to measure the height of the test specimen. (Jungang Liu, 2019).116
4.8 Looking down on the base of the triaxial cell. (Jungang Liu, 2019)...................117
4.9 Rubber membrane placed on the base of the triaxial cell. (Jungang Liu, 2019).....117
4.10 Cylindrical mold placed on the base of the triaxial cell. (Jungang Liu, 2019)...118
4.11 The membrane adhereon the inner wall of the mold. (Jungang Liu, 2019)...........118
4.12 Soil sample placed within the zero raining device. (Jungang Liu, 2019)..........119
4.13 Porous stone placed on top of the Monterey No. 0 Sand Specimen. (Jungang Liu, 2019). 119
4.14 Top cap placed on top of the specimen (Jungang Liu, 2019)...........................120
4.15 Positioning of the top surface of the loading cap parallel to the base of the triaxial cell
(Jungang Liu, 2019)..............................................................120
4.16 Measuring the length of the sample using a vernier caliper (Jungang Liu, 2019)......121
4.17 Pi tape (Jungang Liu, 2019).....................................................122
4.18 Specimen was saturating (Jungang Liu, 2019).........................................122
4.19 Set up on the computer to run MTS. (Jungang Liu, 2019)..............................127


4.20 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 30% and effective stress 15psi.
(Jungang Liu, 2019)....................................................................130
4.21 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with
Different percent of fines contents at relative density of 30% and effective stress 30psi. (Jungang Liu, 2019)................................................................131
4.22 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with
Different percent of fines contents at relative density of 45% and effective stress 15psi. (Jungang Liu, 2019)................................................................131
4.23 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with
Different percent of fines contents at relative density of 45% and effective stress 30psi. (Jungang Liu, 2019)................................................................132
4.24 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with
Different percent of fines contents at relative density of 60% and effective stress 15psi. (Jungang Liu, 2019)................................................................132
4.25 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with
Different percent of fines contents at relative density of 60% and effective stress 30psi. (Jungang Liu, 2019)................................................................133
4.26 Cyclic Axial Load versus Number of cycles to liquefaction in Cyclic Triaxial Test.
(Jungang Liu, 2019)....................................................................134
4.27 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test.
(Jungang Liu, 2019)....................................................................135
4.28 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. (Jungang Liu, 2019).136
4.29 P versus q stress path in Cyclic Traixial Test. (Jungang Liu, 2019).....................137
5.1 Idealized stress and strain components within the HCA subjected to axial load, W, torque, Mr, internal pressure, Pi, and external pressure, Po: (a) hollow cylinder coordinates; (b) Element component stresses; (c) element component strains; (d) element principal stresses
(After Zdravkovic and Jardine, 2001)..................................................140
5.2 Definitions of average stresses and strains (after Hight et al., 1983)...............146
5.3 Definitions used for stress non-uniformity and accuracy (after Hight et al., 1983)....147
5.4 Effect of stress ratio level on non-uniformity coefficients (after Vaid et al., 1990).151
5.5 Shear stress distribution in hollow cylinder torsional shear test specimens (after Porovic,


1995)..................................................................................152
5.6 Mohr’s circle for determining maximum vertical stress in HCTA-88 (Chen, 1988)..........168
5.7 Mohr’s circle for determining maximum shear stress in HCTA-88 (Chen, 1988).............168
5.8 Schematic diagram of the HCTA-88 at the University of Colorado at Denver Geotechnical
Laboratory (Chen, 1988)................................................................169
5.9 Bottom Platen (Jungang Liu, 2019).....................................................170
5.10 Base Plate (Jungang Liu, 2019).......................................................170
5.11 Bottom Sample Pedestal (left) and Top Sample Pedestal (Jungang Liu, 2019)............171
5.12 Top Cap and Top Plate (Jungang Liu, 2019)............................................171
5.13 Piston Locking Mechanism (Jungang Liu, 2019).........................................172
5.14 Supporting Bars (Jungang Liu, 2019)..................................................172
5.15 Pressure Chamber (Jungang Liu, 2019).................................................173
5.16: Hollow cylinder test setup in software of MTS. (Jungang Liu, 2019)...................175
5.17 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test. (Jungang
Liu, 2019)............................................................................177
5.18 Pore water pressure change versus Number of cycles to liquefaction in CHCT. (Jungang
Liu, 2019)............................................................................178
5.19 Cyclic shear stress versus shear strain in Hollow Cylinder Test. (Jungang Liu, 2019).179
5.20 p versus q stress path in Hollow Cylinder Test. (Jungang Liu, 2019)..................180
6.1 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test (Liu, et al,
2017)..............................................................................191
6.2 Cyclic Axial Load versus Number of cycles to liquefaction in Cyclic Triaxial Test. (Liu, et
al, 2017)..............................................................................192
6.3 Pore water pressure change versus Number of cycles to liquefaction in HC Test. (Liu, et al
2017)..............................................................................194
6.4 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. (Liu, etal, 2017).............................................................................195


6.5 Cyclic shear stress versus shear strain in Hollow Cylinder Test (Liu, et al, 2017).....197
6.6 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. (Liu, et al, 2017).198
6.7 p versus q stress path in Hollow Cylinder Test. (Liu, et al, 2017).....................200
6.8 p versus q stress path in Cyclic Traixial Test. (Liu, et al, 2017).....................201
7.1 SPT case histories of cohesionless soils with FC > 35% and the NCEER Workshop (1997)
Curve and the recommended curves for both clean sand and for FC = 35% for M = I'A and o Vo = 1 atm (TM.Idriss and R.W.Boulanger, 2004).................................205
7.2 SPT case histories of cohesionless soils with 5% Both clean sand and for FC = 15% for M = I'A and o Vo = 1 atm (TM.Idriss and R.W.Boulanger, 2004)...........................................................206
7.3: SPT case histories of cohesionless soils with 15% IV2 and o Vo= 1 atm (TM.Idriss and R.W.Boulanger, 2004).................................207
7.4 Stress Ratio Required to reach Liquefaction in 10 Cycles versus Fine Content Under
Confining pressure 15 psi (N.Y. Chang, 1990)............................................208
7.5 Stress Ratio Required to reach Liquefaction in 30 Cycles versus Fine Content Under
Confining pressure 15 psi (N.Y.Chang, 1990).............................................208
7.6 Stress Ratio Required to reach Liquefaction in 100 Cycles versus Fine Content Under
Confining pressure 15 psi (N.Y.Chang, 1990).............................................209
7.7 Cyclic resistance ratio versus equivalent (FC)400 (Tzuo-Shin Ueng, Chia-Wen Sun and
Chieh-When Chen, 2013)..................................................................209
7.8 Variation in Liquefaction Resistance with Fines Content for Ni=20 (Yong Wang and Yanli
Wang, 2010).............................................................................210
7.9 Change of relative density and liquefaction potential with different fines contents and Silts
For tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011).................210
7.10 Change of void ratio and liquefaction potential with different fines contents and silts for
Tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011)....................211
7.11 (a) Sand with low fines content (b) Sand with high fines content (Shaoli Yang, Suzanne
Lacasse and Rolf Sandven, 2005)........................................................213
7.12 Effects of fines on Binary Packing of Spherical Particles: (a) variation in the volume of


Voids and solids (b) variation on emin. (Misko Cubrinovski and Kenji Ishihara, 2002).214
7.13 Variation in emax and emin with fine content of mixtures of Cambria Sand and Nevada fines.
(Misko Cubrinovski and Kenji Ishihara, 2002)..........................................215
7.14 Intergranular soil mix classification. (S. Thevanayagam; T. Shenthan; S. Mohan and J.
Liang, 2002)...........................................................................216
7.15 Stress ratio require for liquefaction in 10 and 35 cycles versus variable fine content (M-L
Mixing samples at relative density 30% and confining pressure 15,30 psi) in cyclic triaxial Tests. (Jungang Liu, 2019).......................................................219
7.16 Stress ratio require for liquefaction in 35 and 50 cycles versus variable fine content (M-L
Mixing samples at relative density 45% and confining pressure 15,30 psi) in cyclic triaxial Tests. (Jungang Liu, 2019).......................................................219
7.17 Stress ratio require for liquefaction in 50 and 90 cycles versus variable fine content (M-L
Mixing samples at relative density 60% and confining pressure 15,30 psi) in cyclic triaxial Tests. (Jungang Liu, 2019).......................................................220
7.18 Stress ratio require for liquefaction in 8 cycles versus variable fine content (mixing samples
At relative density 30% and confining pressure 15psi) in cyclic hollow cylinder tests. (Jungang Liu, 2019)..............................................................221
7.19 Stress ratio require for liquefaction in 27 cycles versus variable fine content (mixing
Samples at relative density 30% and confining pressure 30 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019).......................................................222
7.20 Stress ratio require for liquefaction in 20 cycles versus variable fine content (mixing
Samples at relative density 60 % and confining pressure 15 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019).......................................................223
7.21 Stress ratio require for liquefaction in 40 cycles versus variable fine content (mixing
Samples at relative density 60 % and confining pressure 30 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019).......................................................224
8.1. Correlation to estimate parameter F in Vucetic-Dobry PWP generation model. (Vucetic and
Dobry, 1986)...........................................................................227
8.2. Excess Pore Water Pressure ratio vs. Volumetric Strain (Chung, R.M., Yokel, F.Y. and
Wechsler, H. 1984).....................................................................229
8.3 Excess pore water pressure build up Test data from NRC (1985).........................230
8.4 The normalized excess pore water pressure ratio versus the normalized number of cyclic Stress cycle with different percent of fines content and relative densities in cyclic triaxial test.


a) relative density at 30% and different of fines content (0%, 5%,10%,15%,25%,35% and 45%). (Jungang Liu, 2019)....................................................231
8.5 The normalized excess pore water pressure ratio versus the normalized number of cyclic Stress cycle with different percent of fines content and relative densities in cyclic hollow Cylinder test, a) relative density at 30% and different of fines content (0%,
5%,10%,15%,25% and 35%).(Jungang Liu, 2019)............................................233
8.6 Average excess pore pressure versus normalized number of cyclic stress cycles in all
Samples a) in cyclic triaxial test with different of fines content (0%, 5%, 10%, 15%, 25%, 35% and 45%).(Jungang Liu, 2019).............................................235
8.7 Curved Equi-Strain Line (Horita, 1985)................................................244
8.8 Flow Chart for Simulation on Strain-Controlled Undrained Cyclic Traixial Tests (Horita,
1985)............................................................................245
8.9 Flow Chart for Simulation on Stress-Controlled Undrained Cyclic Triaxial Tests (Horita,
1985)............................................................................246
8.10- 1 Comparison between Measured and Simulated Responses of Soil Samples in Generation
of Pore Water Pressure. (Jungang Liu, 2019)..........................................247
8.10- 1 Comparison between Measured and Simulated Responses of Soil Samples in Generation
of Pore Water Pressure. (Jungang Liu, 2019)....................................248
8.11 Comparison between Measured and Simulated Responses of Saturated Monterey No. 0/30 Sand to Stress-Controlled Undrained Cyclic Loading with stress ratio = 0.4, No. cycles to Liquefaction =10 a) Effective Stress Path (Jungang Liu, 2019)........................249
10.1 SPT case histories of cohesionless soils with FC > 35% and the NCEER Workshop (1997)
Curve and the recommended curves for both clean sand and for FC = 35% for M = I'A and o Vo = 1 atm (I.M.Idriss and R.W.Boulanger, 2004)...........................267
10.2 Shear stress reduction factor, rd, relationship (I.M.Idriss and R.W.Boulanger, 2010).268
10.3. Overburden correction factor (CN) relationship for CPT and SPT penetration resistances:
(a) for Oy /Pa = 0 - 10, and (b) for ov/Pa = 0-2 along with Liao and Whitman's (1986) Relationship................................................................269
10.4. Case Histories from Idriss and Boulanger (all data for soil liquefy): (Ni)6o versus FC (%)
Was created by Jungang Liu (2019)...............................................271
10.5. Case Histories from Idriss & Boulanger: Average (Nl)60 versus FC (0%-25%) was created By Jungang Liu (2019).................................................................272


10.6. Case Histories from Idriss & Boulanger: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average (Nl)60 with different FC (0%-20%) was created by Jungang Liu (2019)............................................................................273
10.7 Case Histories from Tokimatsu &Yoshimi (all data for soil liquefy): N1 versus FC (%) was
Created by Jungang Liu (2019)...................................................279
10.8 Case Histories from Tokimatsu &Yoshimi: Average (Nl), average-MM (Nl) versus FC
(0%- 20%) was created by Jungang Liu (2019).....................................280
10.9. Case Histories from Tokimatsu &Yoshimi: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average Nrwith different FC (0%-20%) was created by Jungang Liu (2019)................................................................................281
10.10.SPT blow count with different fines content (0%-25%) in case histories data from Idriss & Boulanger and Tokimatsu &Yoshimi (all data for soil liquefy) created by Jungang Liu (2019)..........................................................................282
10.11 Relationship between void ratio range and fines content of sandy soils (Misko.
Cubrinovshi & Kenji. Ishihara, 2002)............................................287
10.11 Cyclic Triaxial test results for calculating (Nl)60 From FC=0%-35%, prepared at Dr=
30%, 45% and 60%. (Jungang Liu, 2019)...........................................292
10.12. Cyclic Triaxial test results for calculating (Nl)60 From FC=0%-35%, prepared at Dr=
30% (Jungang Liu, 2019)...........................................................293
10.13 Cyclic Triaxial test results for calculating (Nl)60 From FC=5%-35%, prepared at Dr=
45% (Jungang Liu, 2019)............................................................294
10.14 Cyclic Triaxial test results for calculating (Nl)60 From FC=0%-35%, prepared at Dr=
60% (Jungang Liu, 2019)............................................................295
10.15 Calculating (Ni)6o versus cyclic stress ratio from cyclic traixial test on relative density of 30% under consolidation pressure 15psi with different fines content. (Jungang Liu,
2019)..............................................................................296
10.16 Calculating (Ni)6o versus cyclic stress ratio in cyclic triaxial test on the relative density of 30% under consolidation pressure of 30psi with different fines content (Jungang Liu,
2019)..............................................................................297
10.17 Calculating (Ni)6o versus cyclic stress ratio in cyclic triaxial test on the relative density of 45% under consolidation pressure of 15 psi with different fines content (Jungang Liu, 2019)......................................................................................298


10.18 Calculating (Ni)6o versus cyclic stress ratio in cyclic triaxial test on the relative density of 45% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019)......................................................................................299
10.19 Calculating (Ni)6o versus cyclic stress ratio in cyclic triaxial test on the relative density of
60% under consolidation pressure of 15 psi with different fines content (Jungang Liu, 2019)......................................................................300
10.20 Calculating (Ni)6o versus cyclic stress ratio in cyclic triaxial test on the relative density of
60% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019)......................................................................301
10.21 Cyclic Hollow Cylinder test results for calculating (Nl)60 From FC=0%-35%, prepared at
Dr= 30% (Jungang Liu, 2019)......................................................303
10.22 Cyclic Hollow Cylinder test results for calculating (Nl)60 From FC = 0%-35%, prepared
At Dr= 60% (Jungang Liu, 2019)...................................................304
10.23 Cyclic Hollow Cylinder test results for calculating (Nl)60 From FC=0%-35%, prepared
At Dr= 30% and 60% (Jungang Liu, 2019)...........................................305
10.24 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (Nl)60 From
FC=0%-35%, prepared at Dr= 30% (Jungang Liu, 2019)...........................306
10.25 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (Nl)60 From
FC=0%-35%, prepared at Dr= 60% (Jungang Liu, 2019)...........................307
10.26 Calculating (Ni)6o from cyclic triaxial and cyclic hollow cylinder tests on relative density 30% & 60% and consolidation pressure 15psi & 30psi, (Ni)6o from case histories versus Different fines content (Jungang Liu, 2019)................................................309


CHAPTER I
INTRODUCTION Problem Statement
Soil liquefaction is a phenomenon is which soil loses much of its strength or stiffness for a short time. Nevertheless it is long enough for liquefaction to be the cause of many failures, deaths and major financial losses. When shaken by earthquakes, a saturated loose granular soil will tend to density. Since the soil is saturated, the densification requires that water be expelled out of the soil mass so that soil particles can become more densely packed. If the process of expelling water cannot occur immediately, the soil particles will tend to become waterborne and induce a rapid rise in pore water pressure. When the induced pore water pressure becomes high enough to counterbalance the total stress acting on soil particles, the granular soil will lose all its shear strength and is, in a broad sense, designated as “in a state of liquefaction failure.”
The enormous damage due to soil liquefaction observed in both in Anchorage, Alaska and Niigata, Japan earthquake of 1964 stimulate geotechnical engineering studies of earthquake-induced liquefaction.
Soil liquefaction is one of the most detrimental forms of earthquake-induced ground failure that can result in catastrophic damage to engineering structures. Ground liquefaction failure usually takes place in saturated loose granular soils. For the past 50 years, liquefaction problems have received a lot of attention and great efforts have been made to understand the liquefaction mechanism and phenomena. However, due to the complexity of liquefaction and spreading problems, such as soil properties (fine content) effect soil liquefaction resistance, limitations on experimental and constitutive model, interaction between liquefiable soil and foundation, and so
l


on, reliable and accurate predictive methods have yet to be developed. Therefore, nowadays, soil liquefaction is still one of the hottest research topics in the geotechnical engineering community.
Objectives and Scope
The major objective of this study is to investigate methods for liquefaction resistance evaluation, the factors affecting liquefaction resistance, investigate the effect of fines inclusion on liquefaction resistance under cyclic triaxial and cyclic hollow cylinder tests, excess pore water pressure generation characteristics and refine current liquefaction resistance assessment procedure and criterion to cover the effect of fines. To achieve the objectives, cyclic triaxial tests and hollow cylinder tests on samples with different percentages of fines of varying fines contents were and will be performed at the Center for Geotechnical Engineering Science at the University of Colorado Denver. Statistical analysis and modeling of test results will be conducted, including the following tasks: 1) review existing methods for liquefaction potential analysis. 2) review factors affecting liquefaction resistance of soils. 3) perform cyclic triaxial and cyclic hollow cylinder tests to evaluate the extent of effects of fines content and plasticity on liquefaction resistance of soils. 4) evaluate relationship of soil liquefaction resistance from between cyclic triaxial and hollow cylinder tests on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content. 5) formulation of statistical relationship for predicting liquefaction resistance of soil containing fines. 6) evaluate threshold fines content effect on liquefiable soil. 7) investion excess pore pressure generation from laboratory test results and HORITA’s constitutive model. 8) refinement of current procedure for evaluating liquefaction resistance of soil covering fines.
2


Significance of This Research
In this research, by using cyclic triaxial and cyclic hollow cylinder tests for evaluating liquefaction resistance, it helps us better understand the relationship between cyclic triaxial and hollow cylinder tests on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content. This research can accurate threshold fines content effect on soil liquefaction resistance, and develop appropriate statistical model for predicting liquefaction resistance of soil containing fines. The new findings of this study can be applied to better realize the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils.
3


CHAPTER II
LITERATURE REVIEW 1:LIQUEFACTION RESISTANCE Liquefaction Resistance from Earthquake Case History Data 1964 Niigata Earthquake
In 1964 at 1:01 p.m. on June 16, 1964, a violent earthquake hit Niigata and Yamagata prefectures, inflicting considerable damage on the city of Niigata for a magnitude 7.5 earthquake. In Niigata, where sand deposits in lowland areas are widespread, the damage was primarily associated with the liquefaction of loose sand deposits. Buildings not imbedded deep in firm strata sank or tilted toward the direction of the center of gravity. Underground structures, such as septic and storage tanks, sewage conduits and manholes, floated up a meter or two above ground level. In flat fields, sand flows and mud volcanos ejected water and sand 2 to 3 minutes after the quake. Liquefied sand 20 to 30 cm thick covered the entire city, as if the whole area had been devastated by flooding. Damage to modem bridges was also extensive. Most notable was the crumbling of five girders of the Showa Bridge, which crosses the Shinano River in the downtown area. The foundation piles were bent excessively due to the loss of lateral resistance of the riverbed sand deposit, and this caused the simply-supported girders to fall. Figure 2.1 is a map of Niigata city, in which the places where liquefaction developed are indicated. It is to be noted that the whole city was built on layers of sand as deep as about 100 meters, although the origins of these deposits are somewhat different from one location to another. A typical shallow-depth soil profile is shown in Fig.2.2 and describes the soil condition at Kawagishicho, where the land was formed about 40 years ago by reclaiming an old river course.
4


Fig. 2.1 Niigata city destroyed by the Niigata Earthquake 1964. (National Geophysical Data Center)
5


Fig-2.2 Soil profile at Kawagishi-cho (Niigata Prefecture) (National Geophysical Data Center) 1964 Alaska Earthquake USA
On Friday, March 27, 1964, a great earthquake of magnitude 9.2 struck Prince William Sound and caused severe damage in the form of landslides and liquefaction.
The major effects of liquefaction in the Alaska earthquake were massive landslides in the cities of Anchorage, Seward, Valdez and around the borders of Kenai Lake. Liquefaction in sand layers, and in sand and silt seams in the clayey soils beneath Anchorage, caused many of the destructive landslides that occurred during the earthquake. The liquefied seams and lenses disturbed the sensitive clays, and caused their strengths to drop below the levels needed for stability. It is shown for an explanation of road embankment failure in Figure 2.3. Lateral
6


spreading in the soil beneath the roadway embankment caused the embankment to be pulled apart, producing the large crack down the center of the road shown in Figure 2.4.
Figure 2.4 Road Embankment Failure Caused by Soil Liquefaction. (Timothy J. Walsh, et. al 1995)
1995 Hyogoken-Nanbu Earthquake
It occurred on Tuesday, January 17, 1995, at 05:46 JST (January 16 at 20:46 UTC) in the southern part of Hyogo Prefecture, Japan. It measured 6.8 on the moment magnitude scale (USGS), and Mj7.3 (adjusted from 7.2) on JMA magnitude scale. The tremors lasted for approximately 20 seconds. The focus of the earthquake was located 16 km beneath its epicenter,
Fig2.3 Cracked Highway in 1964 Alaska Earthquake (Timothy J. Walsh, et. al 1995)
7


on the northern end of Awaji Island, 20 km away from the city of Kobe. Figure 2.5 shows the locations of the aftershocks.
The liquefaction and non-liquefaction data that were used in the subsequent liquefaction triggering analysis (Moss, 2003) are summarized in the accompanying Table 2.1. Figure 2.6 is a map of the greater Kobe Port region that shows the approximate locations of these CPT Field case history sites (which have a three letter designation corresponding to the trace); superimposed over SPT field case history sites (which have number designation corresponding to the log) and PGA contours, the latter two are from Cetin et al. (2000)
oo
1 Jnuin-i PrcMTiwcxi KobCJu.il limltulr. Kjotn I'nivcnm
t ^nrrnWr- »■ >rr» not* rfl*’ PfTVfM* >n R* iTiKiti fanmFr At* f Wvcmr
KrwwH Mmw f(i tatw htc**' of Smncr iLirhi Ltoinr\w
Fig 2.5 Location of aftershocks (Moss, 2003)
8


Table 2.1: Summary of CPT Field case histories from the 1995 Hyogoken-Nambu (Kobe) earthquake. (Moss, 2003)
I EVENT | M, ±|
IMSHyngofcen-Nanbu 7.20 0.1)
SITE [DESCRIPTION LIQ? DATA CLASS CR1T LAYER ± c o,' (KPa) ± Qc,l,mod (MPa) CSR*
Dusl Management Center V B 6 to 8 7.00 0.33 2.00 0.31 0.11 7.83 2.53 0.49 0.20 0.64 70.45 4.92 7.83 0.29
Ima2u Elementary School V C 8.0 to 12.0 10.00 0.67 1.40 0.40 0.17 0.80 0.19 0.80 0.34 0.90 101.43 7.23 1.14 0.37
Koyo Junior High School V B 6.5 to 7.5 7.00 0.17 4.00 0.28 0.10 8.03 0.54 1.24 0.87 0.50 95.07 3.96 8.76 0.27
Kobe Customs Maya Office A Y B 4 to 9 6.50 0.17 1.80 0.45 0.16 2.93 0.34 0.40 0.13 0.78 75.24 3.97 2.93 0.43
Kobe Customs Maya Office B Y B 3 to 6 4.50 0.17 1.80 0.48 0.15 6.98 0.73 0.87 0.17 0.54 55.86 3.12 7.41 0.45
Kobe Pori Const. Office Y B 3 to 5 4.00 0.13 2.50 0.42 0.13 5.99 1.15 0.29 0.11 0.76 55.79 2.91 5.99 0.40
Koyo Pump Station Y B 5 to 6 5.50 0.17 2.60 0.33 0.11 238 0.57 1.75 0.82 0.65 71.00 3.41 3.68 0.31
Kobe Wharf Public Co. Y B 4.0 to 5.5 4.75 0.25 2.10 0.35 0.12 6.03 0.74 0.78 0.40 0.65 60.33 3.41 6.33 0.33
Koyo Elementary School Y B 6.510 7.0 6.75 0.17 4.20 0.28 0.10 2.93 1.44 2.17 1.50 0.54 94.01 3.91 4.56 0.26
Mlzukasa Park Y C 6.9 to 7.9 7.40 0.17 2.00 0.45 0.16 1.63 0.60 0.99 0.48 0.75 85.33 4.36 2.19 0.43
Shiporex Kogyo Osaka Factory Y B 4.0 to 7.0 5.50 0.33 1.50 0.37 0.12 3.93 2.18 0.41 0.24 0.74 54.71 4.44 3.93 0.34
Hamakoshlenn Housing Area Y B 2.5 to 5.0 3.75 0.42 200 0.38 0.13 7.00 1.51 0.65 0.22 0.59 49.96 3.85 7.16 0.36
Talto Kobe Factory Y B 3.2 to 4.2 3.70 0.17 1.60 0.39 0.13 4.85 0.86 0.39 0.12 0.75 42.13 3.36 4.85 0.36
Tokuyama Concrete Factory Y B 4.0 to 4.8 4.40 0.13 2.00 0.40 0.13 2.55 0.88 0.40 0.19 0.80 50.98 3.48 2.55 0.38
Nisseki Kobe Oil Tank A Y B 4.8 to 6.1 5.45 0.22 2.40 0.43 0.14 5.30 1.31 0.61 0.36 0.72 69.15 3.53 5.42 0.41
Ntesekl Kobe Oil Tank B Y B 5.0 to 6.0 5.50 0.17 2.40 0.43 0.14 6.25 1.34 0.74 0.27 0.70 69.64 342 6.53 0.41
New Port No. 6 Pier Y B 3.5 to 5.5 4.00 0.33 250 0.42 0.14 9.47 1.60 0.43 0.11 0.70 55.79 3.55 9.47 0.40
Mlnatojima Junior Hfcgh Y B 4.0 to 4.5 4.25 0.08 2.70 0.32 0.10 4.71 1.35 0.94 0.42 0.65 59.57 2.91 5.16 0.30
New Wharf Const-Offices Y B 3.2 to 3.8 3.50 0.10 2.60 0.31 0.10 3.56 0.81 0.93 0.64 0.64 51.62 2.76 4.00 0.29
Fukuzuml Park N C 11.0IO 12.5 11.75 0.33 3.10 0.36 0.18 17.09 3.45 1.42 0.57 0.40 115.94 6.85 18.06 0.33
Honjyo Central Park N B 4.0 to 6.0 5.00 0.33 2.50 0.48 0.16 17.30 3.75 0.60 0.25 0.56 70.48 3.96 17.42 0.45
Kobe Art Institute N B 3.5 to 3.8 3.65 0.05 3.00 0.32 0.10 13.64 5.38 1.90 1.31 0.33 57.62 2.66 15.08 0.30
Vos hi da Kogyo Factory N B 3 to 5 4.00 0.33 3.00 0.33 0.11 9.43 7.22 2.71 2.73 0.34 59.19 3.64 11.72 0.31
Shimonakajima Park N B 3.0 to 4.5 3.75 0.17 2.00 0.50 0.16 19.49 0.80 0.73 0.43 0.53 46.11 3.38 19.77 0.47
SumlyosN Elementary N B 2.4 to 3.2 2.80 0.13 1.90 0.43 0.14 17.35 4.20 0.66 0.31 0.54 38.09 3.15 17.53 0.41
Nagashi Park N B 1.1 to 1.8 1.45 0.12 1.00 0.49 0.16 14.51 4.31 1.05 0.49 0.51 21.59 2.32 15.16 0.46
A
•«ip
—SttP
2
â–  Sebimograph U non Liquefied * L i quefi td
Fig 2.6 Map Showing CPT Field Case History Locations from the 1995 Hyogoken-Nambu Earthquake, in relation to the SPT Case History Locations and PGA Contours from Cetin et al.
(2000).
9


1999 Chi-Chi Earthquake (Taiwan)
At 1:47am on the morning of September 21, 1999, the largest earthquake of Taiwan’s recent history hit central Taiwan at 7.3 magnitude causing thousands of deaths, building collapses, destruction to bridges, highways, dams, and railways. Schools were closed, power was cut, and people were evacuated.
The Chi-Chi Earthquake caused widespread damage to all areas of Taiwan, particularly in Yuanlin, Nantou, Wufeng, Chang-Bin, as well as the Port of Taichung. Most of the damages due to the earthquake were from ground failure mechanisms such as liquefaction and landslides.
Ku, Chih-Sheng, Der-Her Lee, and Jian-Hong Wu (2003) showed that two hundred and seventy five cone penetration test data were collected from the liquefaction-affected areas, and 46 liquefaction case histories and 88 non-liquefaction case histories were derived that can be used to evaluate the accuracy of existing liquefaction evaluation models. In Figure 2.7 shows the conventional plot of CSR7.5 versus qcm for the 134 cases. It showed that almost all liquefied cases have a qcwv alue of less than 50, which might be result of disturbance due to liquefaction. To develop the “unbiased” boundary curve, the qcm values of most if not all liquefied cases must be increased proportionally. However, no guidance is available for such correction in the liquefied cases. An alternative is to give more “weight” to the non-liquefied cases. In other words, the boundary curve is drawn not to encompass the liquefied data points, but rather to exclude all non-liquefied cases. Taking this approach, a boundary curve is obtained and shown in Figure 2.7. Symbolically,
CRR7.5 = 0.058 exp [0.02 qcm\
10


tn
(4
ce
CO
o
0 50 1 00 150 200 250
Normalized Cone Resistance, qc-|N
Figure 2.7 Proposed CPT-based liquefaction boundary curve (Ku, Chih-Sheng, Der-Her Lee, and Jian-Hong Wu 2003)
2011 Tohoku Earthquake
A devastating earthquake of moment magnitude Mw 9.0 struck the Tohoku and Kanto regions of Japan on 11th March, 2011 at 2:46 pm. The earthquake caused great economic loss, loss of lives and tremendous damage to structures and infrastructures. This was the largest earthquake ever recorded in Japan and one of the five most powerful earthquakes in the world since modem record keeping began in 1900.
Widespread liquefaction was observed in certain areas with plenty of evidence of sand boils. Fig.2.8 shows photographs of sand boils on the morning of 12th March. While Fig 2.5(left) shows the sand boils in the Takasu Park, Fig. 2.8(right) shows the liquefaction in the paved car
ll


park of the Disneyland amusement park. The liquefied and ejected soil consisted of different types of materials ranging from pure brown sand with small fines content to grey silty sand, and also in some locations dredged recycled material.
Fig 2.8 Observed liquefaction in the park (left) and parking area of Disneyland (right). (S. Bhattacharya, et, 2011)
For Fig 2.9, a simplified stress-based liquefaction triggering analysis was carried out by considering a typical soil profile in Urayasu, where widespread liquefaction and sand boils were observed. The calculated probability of liquefaction over the depth up to 20 m is presented in the third sub-figure in Fig. 2.9. A probabilistic model of liquefaction, developed by Cetin et al., based on the database of field performance case histories. However, the probabilistic model of liquefaction has not been validated/calibrated against a mega thrust Mw 9.0 subduction event. Bhattacharya, et al (2011) indicated that soil liquefaction is highly likely to trigger at shallow depths (up to 6 m) and medium depth (12-16m), and the effects of the moment magnitude are not so significant at these depths in the Figure 2.9. The main reason for the high liquefaction potential for this profile is the low blow count, N (less than 5).
12


0
CL
Q
2
4
6
8
10
12
14
16
18
20
â–¼
Sandy silt -f
Silty fine sand °v
Sandy silt
Fine sand
Sandy silt
Soil type
0 10 20 30
Ncount
Liquefaction
probability
Max = 8.5 Max /Ww = 8.0 Max A/fw = 9.0
Input parameter:
PGA: 142.9 cm/s2 \/si2- 123 m/s Water table: 1 m Dry density: 1.76 g/cm3 Wet Density: 1.92 g/cm3
Fig. 2.9 Typical soil profile in Urayasu (S. Bhattacharya, et, 2011)
2011 Christchurch Earthquake New Zealand
The 2011 Christchurch (Canterbury) Earthquake was a magnitude 6.3 event that occurred at 12:15 p.m. local time on 22 February on the South Island of New Zealand. This was a shallow earthquake and its epicenter distance to the east side of Christchurch is less than 6 kilometers. There was extensive liquefaction and ground failure, which led to widespread loss of water and electricity. The earthquake produced strong ground motions within the central business district (CBD) of Christchurch
Much of the Christchurch region is situated on unstable liquefiable sandy soil. The poor soil condition caused damage during the 2011 earthquake (see Figure 2.10). As shown in the figure, liquefaction (sand boils at the surface) was evident in many locations. The ground failure was extensive and affected residential buildings as well as a major roadway. For example, ground failure damaged roads in this area, rendering many of them inoperable. Figure 2.10 also
13


shows damage to a bridge abutment. The abutment cracked as the soil in the approach slab failed. Christchurch has several rivers, including the Avon and Heathcote Rivers which contributed to the underlying soil problems. Also exacerbating the problem was the fact that many water mains ruptured, pouring water onto the underlying soil and resulting in damage.
The resulting liquefaction documentation map is shown in Figure 2.11. Three areas of different liquefaction severity are indicated: A) moderate to severe liquefaction (red zone), B) low to moderate liquefaction (yellow zone), and C) liquefaction predominantly on roads with some on properties (pink zone).
Preliminary geotechnical zoning based on existing data indicates several different areas within the CBD that are dominated either by gravelly layers, thick liquefiable sands or sandy silt mixtures, and peat in the top 8-10 m of the deposits. The soil profiles and thicknesses of these layers are highly variable even within a single zone, thus imposing difficult foundation conditions and sometimes resulting in unconventional or hybrid types of foundations being adopted for buildings. The gravelly soils, even though relatively more competent foundation soils, typically show medium standard penetration test (SPT) N values of about 15 to 25 blow counts, whereas the liquefiable loose sands and silt-sand mixtures have low resistance of less than N = 12 or cone penetration test (CPT) qc values less than 3-6 MPa.
Major earthquake case histories are shown in Table 2.2. It showed that many earthquakes caused ground failure and damaged engineering structures.
14


Figure 2.11 Liquefaction documentation map of eastern Christchurch from drive-through Reconnaissance. (Christchurch, New Zealand Earthquake Field Report, 2011)
Soil liquefaction sand boll
Land sltdu
Broken water main Flooded roadway
Figure 2.10. Ground Failures and Resulting Water Damage. (Christchurch, New Zealand
Ground failure under roadway
Abutment sliding
15


Table 2.2 Major earthquake case histories
Earthquake Remarks Post-earthquake developments
1908 Reggio Messina earthquake (Italy) 120,000 fatalities. A committee of nine practicing engineers and five professors were appointed by Italian government to study the failures and to set design guidelines. Base shear equation evolved i.e. the lateral force exerted on the structure is some percentage of the dead weight of the structure. (Typically 5-15%)
1923 Kanto earthquake (Japan) Destruction of bridges, buildings. Foundations settled, titled and moved. Seismic coefficient method (equivalent static force method using a seismic coefficient of 0.1-0.3) was first incorporated in design of highway bridges in Japan (MI 1927)
1933 Long Beach earthquake (USA) Destruction of building specially school buildings UBC (1927) revised. This is the first earthquake for which acceleration records were obtained from the recently developed strong motion accelerograph.
1964 Niigata earthquake (Japan) Soil can also be a major contributor of damage. Soil liquefaction studies started.
1971 San Fernando earthquake (USA) Bridges collapsed, dams failed causing flood. More soil effects were observed. Liquefaction studies intensified. Bridge retrofit studies started.
1976 Tangshan earthquake (China) Most of the liquefaction occurred in the upper few meters of loose to medium-dense silty fine sand or fin-to medium-clean sand More research on soil liquefaction. Using SPT to investigate soil liquefaction.
1994 Northridge earthquake (USA) Steel connections failed in bridges Importance of ductility in construction realized. Significant damage potential due to near-fault motions was recognized.
16


1995 Kobe earthquake (Japan) Massive foundation failure. Soil effect was the main cause of failure. Downward movement of a slope (lateral spreading) is said to be one of the main causes of foundation failure. JRA (1996) code modified (based on lateral spreading mechanism) for design of bridge piles.
1999 Chi-Chi earthquake (Taiwan) Many bridges collapsed as they were located close to the faults Special care must be undertaken while designing important structure in the vicinity of the plate boundaries and faults
1999 Koceli earthquake (Turkey) Damage to Bolu tunnel due to fault movement. Damage to buildings and bridges. However, buildings conforming with the design codes performed well. Research on fault induced failures intensified.
2001 Bhuj earthquake (India) Large-scale destruction. Good performance of some new jetties of the Kandla port. Tilting of the Kandla Tower building without any damage. Large diameter concrete piles performed better than small diameter ones. Steel piles with filled in concrete also performed better.
2004 Sumatra earthquake and tsunami Destruction to build environment due to earthquake and giant tsunami waves. New research focused on tsunami warning systems
2011 Tohoku earthquake and tsunami The most powerful earthquake ever recorded to have hit Japan, and the fourth most powerful earthquake in the world. The tsunami caused nuclear accidents. More research focused on earthquake and tsunami cause nuclear power plant.
17


Field Methodologies for Liquefaction Potential Evaluation
General
Preliminary assessments may often be made to determine whether a given site is likely or not likely to liquefy in response to earthquake ground motions. The previous occurrence of liquefaction in site soils, knowledge of embankment placement techniques that have historically performed well or poorly when shaken, the seismicity of the site, and degree of saturation are some of the factors that may indicate the potential for future liquefaction.
The importance of adequate site characterization to seismic stability analysis cannot be overstated. Much can and should be accomplished by acquiring and examining existing site data from the geological literature, historical records, earlier field investigations and even sensing imagery before additional subsurface investigation is planned or undertaken. The following information is essential to initial assessment of the potential for earthquake-induced ground failure:
(1) Site topography
(2) Soil profile, including general classification soil properties and the origin of site soils
(3) Water level records, representative of both current and historical fluctuations
(4) Evidence from project records, aerial photographs, or previous investigations of past ground failure at the site or at geologically and seismological similar areas (including historical records of liquefaction, topographical evidence of landslides, sand boils, effects of ground movement on trees and other vegetation, subsidence, and sand intrusions in the subsurface)
(5) Seismic history of the site
18


(6) Geologic history of the site, including age of site soils, glacial preconsolidation or preconsolidation by now-eroded overburden, and lateral extent and continuity of soil deposits.
A subsurface investigation should be performed in two phases, distinguished by coverage and purpose. The first of these should include Standard Penetration Tests (SPT) for measuring penetration resistance and obtaining disturbed split-spoon samples for classification and water content determination. Coverage of the site with SPT borings should be adequate to (1) establish general soil conditions, distributions of soil types, homogeneity and ground water elevations; (2) identify soils that, if shaking were sufficiently intense, might liquefy; and (3) assist in specifying the locations of additional boring and geophysical surveys aimed at detailed seismic response evaluation. The second phase of subsurface investigation likely includes surveys and undisturbed sampling borings to: (1) refine preliminary interpretation of stratigraphy and the extent of potentially liquefiable soils; (2) measure in situ densities and dynamic properties for input to dynamic response analysis; (3) recover undisturbed soil samples for laboratory testing.
Standard Penetration Test (SPT)
General
The Standard Penetration Test (SPT) is a soil-sampling procedure that is in worldwide use and is generally accepted as providing some correlation with in-place properties of a soil. The SPT requires that a 2-in. (51mm) split spoon sampler be used in conjunction with a 140-Ib (63.6kg) drive weight. The SPT reports the number of blows N to drive the sampler 1ft (0.3m) into undisturbed soil by using the 140-Ib weight falling 30in. (0.76m).
19


In the United States and most other countries, the standard penetration test (SPT) has been the most commonly used in situ test for characterization of liquefaction resistance; factors that tend to increase liquefaction resistance (e.g. density, prior seismic straining, over consolidation ratio, lateral earth pressures, and time under sustained pressure) also tend to increase SPT resistance. Seed et al. (1983) compared the corrected SPT resistance and cyclic stress ratio for clean sand (Figure 2.12) and silty sand (Figure 2.13) sites at which liquefaction was or was not observed in earthquakes of M = 7.5 to determine the minimum cyclic stress ratio at which liquefaction could be expected in a clean sand of a given SPT resistance.
Cyclic Resistance Ratio from the Standard Penetration Test
The cyclic resistance ratio represents the liquefaction resistance of the in situ soil. The most commonly used method for determining the liquefaction resistance is to use the data obtained from the standard penetration test. The advantages of using the standard penetration test to evaluate the liquefaction potential are as follows:
1. Groundwater table: a boring must be excavated in order to perform the standard penetration test. The location of the groundwater table can be measured in the borehole. Another option is to install a piezometer in the borehole, which can then be used to monitor the groundwater level over time.
2. Soil type: In clean sand, the SPT sampler may not be able to retain a soil sample. But for most other types of soil, the SPT sampler will be able to retrieve a soil sample.
The soil sample retrieved in the SPT sampler can be used to visually classify the soil and to estimate the percent fines in the soil. In addition, the soil specimen can be returned to the laboratory, and classification tests can be performed to further assess the liquefaction susceptibility of the soil.
20


3. Relationship between N value and liquefaction potential: in general, the factors that increase the liquefaction resistance of a soil will also increase the (Ni)6o from the standard penetration test. For example, a well-graded dense soil that has been preloaded or aged will have a higher resistance to liquefaction and will have high values of (Ni)6o. Likewise, a uniformly graded soil with a loose and segregated soil structure will be more susceptible to liquefaction and will have much lower values of (Ni)60.
The presence of fines can affect SPT resistance and therefore must be accounted for in the evaluation of liquefaction resistance (Seed et al., 1985; Ishihara and Kosecki, 1996; Koester, 1994). Examination of Figures 2.12 and 2.13 shows that the liquefaction resistance of sands is not influenced by fines unless the fines comprise more than 5% of the soil. At higher fines contents, the fines tend to inhibit liquefaction [i.e., the CSR required to initiate liquefaction (for a given (Ni)6o value)]. (Seed et al., 1985; Ishihara and Kosecki, 1996).
However Hsing-Cheng Liu (1992) showed that at a small fine content the increase in fines content leads to the decrease in liquefaction resistance until the fines content reaches a critical value where liquefaction resistance is the lowest. Beyond the critical fines content, the liquefaction resistance start to increase with increasing fines content in laboratory testing.
The plasticity of the fines can also influence liquefaction resistance; the adhesion of plastic fines tends to resist the relative movement of individual soil particles and thereby reduce the generation of excess pore pressure during earthquakes. Laboratory tests (Ishihara and Koseki, 1996) indicate little influence at plasticity indices below 10, and a gradual increase in liquefaction resistance at plasticity indices greater than 10. Ishihara (1996) suggested that the effects of plasticity could be accounted for by multiplying the CSR by the factor “F”
21


1.0
PI < 10
F 1
1.0+ 0.022 (PI - 10) PI> 10
Since most sandy soils in alluvial deposits and man-made fills have plasticity indices less than about 15, the effect of fines plasticity is usually small. Because strong-motion duration (hence equivalent number of uniform stress cycles) increases with earthquake magnitude, the minimum cyclic stress ratio required to initiate liquefaction decreases with increasing magnitude. The minimum cyclic stress ratio for other magnitudes may be obtained by multiplying the cyclic stress ratio for M= 7.5 earthquakes by the factors shown in Table 2.3.
22


Figure 2.12 Relationship between cyclic stress ratios causing liquefaction and (Ni)6o values for clean sands in M= 7.5 earthquakes. (After Seed et al. (1975)
23


0 10 20 30 40 50
(^i)eD
Figure 2.13 Relationship between cyclic stress ratios causing liquefaction and (Ni)6o values for silty sands inM= 7.5 earthquakes. (After Seed et al. (1975).
24


Table 2.3 Magnitude correction factors for cyclic stress approach (Seed et al, 1975)
Magnitude, M CSRm / CSRm= 7.5
5.25 1.50
6 1.32
6.75 1.13
7.5 1.00
8.5 0.89
A large data base of SPT blow counts, normalized to account for the effects of different overburden pressure and performance conditions, has been correlated to occurrence and nonoccurrence of liquefaction in a wide variety of soils (Seed, Idriss and Arango, 1983, Seed, et al. 1985, Farrar, 1988).
The SPT remains the tool of choice for preliminary in situ investigation of liquefaction potential as a result of its empirical correlation to field performance. The term “standard” is of dubious relevance, as the standard procedure specified for SPT performance by the American Society for Testing and Materials (1967) is not so rigid as to prevent variations in practice. Other countries have also developed indigenous versions of the test, unconstrained by the US regulation. Base on the standard penetration test and field performance data, Seed et al. (1985) concluded that there are three approximate potential damage ranges that can be identified show in Table 2.4.
Table 2,4 Relationship between (Ni)6o and potential damage. (Seed et al, 1985)
(Ni)60 Potential damage
0-20 High
20-30 Intermediate
>30 No significant damage
Figure 2.13 presents a chart that can be used to determine the cyclic resistance ratio of the in situ soil. This figure was developed from investigations of numerous sites that had liquefied or did not liquefy during earthquakes. Use Figure 2.13 to determine the cyclic resistance ratio of the in situ soil, as follows: 1. Standard penetration test (Ni)6o value: note in Figure2.13 that the
25


horizontal axis shows data from the SPT test, which must be expressed in terms of the (Ni)6o values. 2. Percent fines: once (Ni)6o value has been calculated, the next step is to determine or estimate the percent fines in the soil. For a given (Ni)6o value, soils with more dines have a higher liquefaction resistance; 3. Cyclic resistance ratio for an anticipated magnitude 7.5 earthquake.
The full, updated database is shown with the Idriss-Boulanger triggering correlation in terms of equivalent CSRm=7.5,o'=i versus equivalent clean sand (Ni)60cs in Figure 2.14. For comparison, the database previously used by Idriss and Boulanger (2004, 2008) is shown in Figure 2.15. Comparison of Figure 2.14with Figure 2.15 indicates that the updated database includes considerably more case histories and that the updated database continues to support the previously derived triggering correlation. Figure 2.16 showed Relationship between cyclic stress ratios causing liquefaction and (Ni)6o values for clean sand with different plasticity index and fine content.
The revised relation for FC< 5% is further compared to other published relations in Figure 2.17, including relations from early in their development (i.e., Seed 1979) to a very recent relation by Cetin et al (2000) that is summarized in Seed et al (2001).
26


0.6 | i i i i | i i i i | i i i i | i i i i | i i
fdriss & Boulanger (2004, 2008) —
O
0.5 ~
s 04
<0
El
b
S o.3
El
S 8
0.2 -
0.1 -
O cP :
O O
o 0 o o
o§ o
o .
AH data
• Liquefaction A Marginal O A/o liquefaction
i i i '---------—----------->
â– 

10
20
(Ni)$ocs
30
40
Figure 2.14 Updated SPT case history database of liquefaction in cohesionless soils with various fines contents in terms of equivalent CSR for M = 7.5 and o'v = 1 atm and equivalent clean sand (Ni)60cs (Idriss-Boulanger, 2008).
27


Cyclic stress ratio
Equivalent clean sand corrected standard penetration, fA/^60cs
Figure 2.15. SPT case history database used previously by Idriss and Boulanger (2004, 2008)
28


(a)
10 20 N
30
40
10
20
30
40
1,60,CS
(b)
N
1,60
Figure 2.16 Parts (a) and (b) of Figure 2.6 in Cetin et al. (2004); note that the points representing case histories are identical in parts (a) and (b) of this figure and as representing conditions with o'v= 1 atm.
29


Cyclic Stress Ratio (CSR)
Modified Standard Penetration - (N^)^q - Blows/ft
Fig 2.17 Curves relating CSR to (Ni)6o published over the past 24 years for clean sands and the recommended curve for M = TA and o'vo = 1 atm (Seed et al 2001).
30


Cone Penetration Test (CPT)
General
The standardized cone-penetration test (CPT) involves pushing a 1.41-inch diameter 55° to 60° cone through the underlying ground at a rate of 1 to 2 cm/sec. CPT soundings can be very effective in site characterization, especially sites with discrete stratigraphic horizons or discontinuous lenses. CPT (ASTM D-3441, adopted in 1974) is a valuable method of assessing subsurface stratigraphy associated with soft materials, discontinuous lenses, organic materials (peat), potentially liquefiable materials (silt, sands and granule gravel) and landslides. The Cone rigs can usually penetrate normally consolidated soils and colluvium, but have also been employed to characterize d weathered Quaternary and Tertiary-age strata. The cone is able to delineate even the smallest (0.64 mm/l/4-inch thick) low strength horizons, easily missed in conventional (small-diameter) sampling programs. Some examples of CPT electronic logs are attached, along with hand-drawn lithological interpretations. Most of the commercially-available CPT rigs operate electronic friction cone and piezocone penetrometers, whose testing procedures are outlined in ASTM D-5778, adopted in 1995. These devices produce a computerized log of tip and sleeve resistance, the ratio between the two, induced pore pressure just behind the cone tip, pore pressure ratio (change in pore pressure divided by measured pressure) and lithological interpretation of each 2 cm interval are continuously logged and printed out.
The CPT is a promising subsurface investigation tool for a variety of applications, particularly in the shallow, soft soils prone to earthquake-induced liquefaction. The electrical friction cone penetrometer (several variations, depending on instrumentation design) have replaced the earlier, mechanical version (both types are sketched in Figure 2.18 and Figure 2.19), due primarily to the ability to obtain continuous, direct measurement of the resistance of
31


soil to penetration and friction. These two parameters have been correlated with soil type and behavioral properties (Douglas, Olsen and Martin, 1994, Olsen and Farr, 1996, and Olsen and Malone, 1994). One failing of the CPT is the inability to obtain physical samples; on the other hand, borings usually require circulation of drilling fluid and much greater labor and time (and accordingly, expense) to advance through similar depths of investigation.
In deference to the large field performance data base on liquefaction potential that has evolved using the SPT, researchers have usually elected to convert CPT data into equivalent SPT values and take advantage of existing correlations. Adaptation of CPT data in this manner is supported by Douglas, Olsen and Martin (1981), who concluded from a detailed study of influential factors common to both the SPT and CPT:
(1) the SPT and CPT are similarly affected by certain soil properties, such that CPT results are directly relatable to SPT results for liquefaction potential;
(2) CPT profiles provide much finer resolution of stratigraphy than do SPT results (and liquefaction failure may occur in thin layers that could lead to sliding); and
(3) The typically large variation of test results associated with the actual performance of an SPT is substantially avoided with the CPT, which is more automated (see also Federal Highway Administration, 1978).
Olsen (2000) proposed normalization of measured CPT data to a function of the effective overburden stress, followed by conversion to continuous, normalized SPT data. The chart shown in Figure 2.20 illustrates the interrelationship developed to predict normalized (to 1 tsf effective overburden stress) SPT blow counts, Ni, the exponent n ranges from about 0.6 in coarse sands to 1.0 in clays (Olsen and Malone, 2000). These normalized SPT data may then be
32


compared to laboratory cyclic strength test data or the field performance data base for various soils.
3 Strain gages
4 Friction sleeve (150 cm2)
5 Adjustment ring
6 Waterproof bushing
7 Cable
8 Connection with rods
Electric Friction-Cone Penetrometer Tip
Figure 2.18 Manufacturing and operating tolerances of cones, taken from ASTM D5778.
33
35.6mm


Example of a Reference Penetrometer With a Fixed Cone and With Friction Sleeve
Figure 2.19 Schematic section through a piezocone head, showing the piezo-element and friction sleeve (from ASTM D5778).
34


iooo
800- 600- 400-
4/1 -
4/3 t- ib 200 -
c |c loose -
r*! 60
1 — ~
II o AO-
o
c
>oJ
_4/l
â– on

cc
a?
d
o
a
“d
O
C3
%_
o
o
t jtjjcc I poor C.r*T (.'H+"t'lii'r4Nj io mc- ^ l.«_f-• J S^T c orr^idl.or=D Lu^cauw O* p GHAVFI
AW
SANDS
Calculated
FU,
N,
based on trends
20-
sa NO
Sand
CLAYS
Predicted
N,
SILT MIXTURES
Sail Chiinu:trriz2luii I■T-Ls^ s.i I n- jl J»ooj canloif
2 â–  SPT:
notary Wash Boring Safety Hammer I wo turns of an old rope
--r----1--■—
O.T 0.2 0.4
Jill—I—J
o.e o.s .
tlW r^ntje rather tliai>
1hr avera^r OPT mcasaeilfierns Oi*( (he * ±t foot «h*p1h mterval when prcdcthHj T
Correeted Friction Ratio (%) in terms ol tsf
FRi il10p = -^l^— ^ q-/roc>f'
too =
U.

^ too
Figure 2.20 CPT prediction of overburden pressure-corrected SPT blow-count (Olsen and Malone, 2000)
35


Cyclic Resistance Ratio from the Cone Penetration Test
As an alternative to using the SPT test, the CPT can be used to determine CRR of the in situ test. The tip resistance from the CPT test can also be used as a measure of liquefaction resistance.
In CPT-based liquefaction evaluations, the tip resistance is normalized to a standard effective overburden pressure of 1 ton/ft2 (96 kPa) by
qci = qc (pa/ Ovo’)0'5 or qci = [1.8/ (0.8+ oV0 )] qc
where oV0 is in tons/ft2 (Kayen et al., 1992). Adjustment for magnitudes other than 7.5 can be made using the CSR correction factors presented in Table 2.3. Kayen et al. (1992) found that liquefaction observations in the 1989 Lorna Prieta earthquake agreed well with the curves of Robertson and Campanella (1985) and Mitchell and Tseng (1990).
For silty sands (> 5% fines), the effects of fines can be estimated by adding the following tip resistance increments to the measured tip resistance to obtain an equivalent clean sand tip resistance (Ishihara, 1993)
The cyclic resistance ratio (CRR) represents the capacity of the soil to resist liquefaction. The relationship recommended by Youd et al. (2001) for computing CRR from CPT measurements can be expressed as (Robertson and Wride 1998):
If (qcm)cs < 50, CRR = 0.833 [((qciN)cs)/1000] + 0.05 If 50 < (qciN)cs < 160, CRR = 93 [((qciN)cS)/1000]3 + 0.08 Where (qciN)cs is the clean-sand cone tip resistance normalized to atmospheric pressure.
The stress-normalized cone tip resistance (qciN)is calculated using the following equation (Robertson and Wride 1998):
qclN = Cq (qc/Pa) = (Pa/Ov )11 (qc/Pa)
36


where qc is the measured cone tip resistance in the same units as Pa, where Pa is a reference pressures assumed to be atmospheric pressure (about 100 kPa) in the same units ov and n is an exponent that depends on soil type. To avoid unreasonably high values at shallow depths, Youd et al. (2001) recommended that Cq be limited to a maximum value of 1.7. For cone measurements made with a pressure transducer behind the cone tip, values of qc are corrected for the effect of pore pressures (Lunne et al. 1997). This correction is particularly significant in silty soils. The exponent n is a variable that depends on soil type and is assumed as 0.5 for granular soils and 1.0 for clay.
Several investigators have noted that liquefaction resistance of soils increases with age (e.g., Seed, 1979; Youd & Hoose, 1977; Youd & Perkins, 1978; Arango et al. 2000; Leon et al. 2006.) However, because the processes causing increased liquefaction resistance with age were poorly understood and proposed correction factors for age had not been verified, Youd et al.
(2001) did not recommended age correction factors at the time of their study. In an effort to account for the affect of age on CRR, the following correction equation has been proposed (Andrus et al. 2004):
CRRa = CRR X Ka2
where CRRa is the age-corrected cyclic resistance ratio, and Ka2 is a factor to correct for influence of age. The value of Ka2 is 1.0 for soils less than a few thousand years old. For older soils, Andrus et al. (2004) suggested using the lower bound of the relationship between cyclic strength and time proposed by Arango et al. (2000).
I.M. Idriss and R.W. Boulanger (2004) showed that the CRR - qcm relation is shown in Figure 2.21 with the case history points for cohesionless soils having FC<5%. The derived relation can be conveniently expressed as:
37


CRR = exp
VcIN
540
+
'IdN
80
1dN
67
+
'IdN
114
This CRR - qciN relation is compared in Figure 2.22 to those by Shibata and Teparaksa (1988), Robertson and Wride (1997), Suzuki et al (1997), and the 5% probability curve by Moss (2003) as summarized in Seed et al 2003.
In liquefaction charts such as Fig. 2.22, every dot represents a saturated sand site which was subjected to a specific earthquake event, and the chart is essentially a way to organize the information available for those sites and their responses during actual earthquakes, which maximizes the chances of predicting if a similar site will or will not liquefy during a future earthquake of the same magnitude (M=7.5). That is, while the way of defining the cyclic stress ratio in Fig. 2.22 and in similar charts was initially suggested by laboratory cyclic test results, both the development and use of a chart like this have their own rules which are unrelated to any cyclic laboratory test. The method relies exclusively on the field measurement of qcm to characterize the liquefaction resistance of the sand.
Table 2.5 Relationship between fines content anc tip resistance increment (Seed et al, 2003)
Fines Content (%) Tip Resistance Increment (tons/ft2)
<5 0
<10 12
<15 22
<35 40
38


Cyclic Stress Ratio, CSR
Normalized Corrected CPT Tip Resistance, qc1N
Fig. 2.21 CPT-based case histories and recommended relation for clean sands for M =71/2 and o'vo = 1 atm (I.M. Idriss and R.W. Boulanger, 2004)
39


Cyclic Stress Ratio, CSR
Fig. 2.22 CPT-based case histories and recommended relation for clean sands with relations proposed by others.
40


Piezometric Cone Penetrometer Test
Miniaturized instrumentation was installed into electric cone penetrometer devices such that pore pressures might be measured both as the probe is pushed into the soil and to monitor ambient head to determine the precise depth to the ground water table. Early studies indicated that penetration of a cone penetrometer would increase pore water pressures in contractive, potentially liquefiable soils and induce negative pore water pressure in dilative deposits (Schmertmann, 1978).
As additional CPT push rod segments are added, excess pore pressure dissipates to ambient levels at the penetrometer tip. Forrest, Ferritto and Wu (1981) report a study of one such device, comparing CPT results (cone tip penetration resistance and piezometric level only) with laboratory cyclic triaxial strengths in waterfront deposits where dissipation rates correlated with permeability; it was postulated that such information could implicate liquefiable soils, though no such attempts were directly made.
Cooper and Franklin (1982) and Norton (1983b) described a piezometer cone penetrometer that measured both cone tip penetration resistance and sleeve frictional resistance as do typical electric CPT devices, with the addition of a pore pressure transducer at the tip
Seismic Cone Penetration Test
Seismic Cone Penetration Testing (SCPT) sounding provides a rapid and cost-effective method for directly measuring shear wave velocity of soils in situ. Shear wave velocity is used as an index of liquefaction resistance since both are influenced by many of the same factors.
41


The seismic cone (Figure 2.23) measures the shear wave velocity of the soil being investigated. Together with a knowledge of the soil saturated unit weight, the shear wave velocity allows an assessment of the small strain shear modulus (Go) and the constrained modulus (Mo) to be made. The small strain shear modulus is an essential input for prediction of ground surface motions from earthquake excitation, evaluation of foundations for vibrating equipment, offshore structures behavior during wave loading, and for prediction of deformations around excavations. The results of the SCPT carried out on Nunavik, Canada are given in Figure 2.25.
The seismic cone is available in 10 and 15 cm2 areas. The cone usually consists of a piezocone unit - measuring the geotechnical parameters qc, fs and U2 - with a receiver for the seismic measurements above it. A schematic diagram, with the layout of the standard technique using a seismic cone, is shown in Figure2.24. The extra equipment needed, in addition to the built in seismometer, is a memory oscilloscope and an impulse source with a trigger for the oscilloscope. The source can consist of a steel beam for shear (S) wave generation or a flat plate for compression (P) wave generation.
The moduli Go and Mo can be determined from the following:
Go = p(Vs)2 (kN/m2)
Mo= p(Vp)2 (kN/m2)
where: p = the soil mass density (kg/m3)
Vs = shear wave velocity (m/sec)
Vp = compression wave velocity (m/sec)
42


SCPT shear wave velocity measurements are used in these evaluations: 1) Liquefaction Risk; 2) Earthquake generated ground-surface movements; 3) Foundations for vibrating equipment; 4)Behavior of offshore structures due to wave loading.
Figure 2.23. Seismic Cone ( from A.P. van den berg)
Figure 2.24. Seismic cone configuration (from A.P. van den berg)
43


Friction Resistivity S wave
ratio (%) (ohm-m) velocity (m/s)
E
f
a
0 10 20 30 40 50 -3 -2 -1 0 1 2 3 1800 2300 2800 3300
Tip load Temperature P wave
(MPa) (°C) velocity (m/s)
Figure 2.25. Results of the SCPT carried out on June 16th 2001 in a permafrost mound near Umiujaq, Nunavik, Canada.
Other Techniques (Appendix A)
44


Laboratory Methods for Liquefaction Resistance Evaluation Undisturbed Sampling
Soil samples are disturbed both mechanically and by change in their effective stress state during sampling and transporting to testing facilities. The term “undisturbed” is liberally interpreted to imply sampling activities that minimize mechanical disturbance for the purposes of this study. As concerns liquefaction potential evaluation, Marcuson and Franklin (1979) review techniques and apparatuses that are still commonly applied to sample granular soils. Significant conclusions reported in that reference include: (1) fixed-piston, thin-walled tube samplers used in boreholes supported by appropriately mixed drilling mud or fluid generally yield high quality samples of many sands; (2) the use of radiographs of samples within sampling tubes permits judgement of sampling disturbance for selection of representative specimens; (3) undisturbed gravel specimens can be successfully obtained only by hand carving larger block samples; and (4) in situ freezing of a larger-than-required volume of soil for subsequent trimming produces very high quality (with regard to mechanical disturbance) soil samples, as long as the freezing front is propagated in a manner that assures free drainage.
Marcuson and Franklin (1979) reported that fixed piston sampling operations tend to produce the best samples so obtained when used in medium dense sands. Tube sampling was observed to densify loose sands and dilate dense sands. The implication is that cyclic strength test results on tube sampled specimens, if interpreted directly, would be unconservative in the case of sands that were loose in situ, and overconservative in dense sands.
Singh, Seed and Chan (2004) examined in situ freezing techniques for undisturbed sampling of saturated sands. Few studies have addressed the efficacy of freezing in silty or
45


clayey soils. Tani and Yasunaka (1988) studied the effects of in situ freezing to sample sands with up to 6% particles finer than 74 micrometers (i.e., passing the US Standard No. 200 sieve). Their results indicated that there was no change in cyclic triaxial liquefaction resistance for alternately frozen and thawed specimens. Samples were taken from the body of a small earth dam that experienced moderate settlement due to liquefaction within either the embankment or its foundation or both. Tani and Yasunaka (1988) claimed from a small number (unstated, but by data plots apparently less than 10) of tests that samples taken by in situ freezing were thus representative of “true liquefaction resistance”.
A number of studies have examined the effects of methods of reconstitution to prepare representative specimens of soils that are difficult to sample (e.g., Mulilis, Chan, and Seed, 1975, Marcuson and Townsend, 1976, Ladd, 1977). No one method of reconstitution best represents natural deposition processes and preserves in situ fabric Laboratory Test Cyclic Triaxial Tests
Ideally, the best cyclic test to evaluate response of soils to earthquake shaking would be one that correctly simulates the loading to which the soil would be subjected in situ. It is commonly believed that at least a single component of earthquake ground motion is adequately reproduced in one form or another of the cyclic simple shear test. Various configurations of cyclic simple shear, cyclic triaxial, large-scale shake table, and cyclic torsional shear (on solid or hollow specimens) apparatuses have been employed to study liquefaction resistance.
In a cyclic triaxial test, a cylindrical specimen of soil encased in a rubber membrane is placed in a chamber, subjected to confining fluid pressure and back pressure, and then loaded axially until failure. The axial load may be applied to the sample through a rigid top platen. The
46


axial force can be compression or extension: thus the axial stress can be either major or minor principal stress. Usually the top platen is laid over a porous stone which allows fluid to flow in and out of the specimen. The axial deformation of the specimen is directly monitored by the movement of the piston which is in contact with or connected to the top platen. The lateral deformation is not usually measured. Transducers are used for pore pressure measurement.
In a cyclic triaxial test, a sample is consolidated under an initial isotropic confining pressure. The confining pressure is kept constant and axial load is either increased (compression test) or decreased (extension test) during a test. Thus, two of three principal stresses are always equal during a test. In a compression test the intermediate principal stress is equal to the minor principal stress; and the axial stress is equal to the major principal stress. In an extension test the major and the intermediate principal stress are equal, while the axial stress is equal to minor principal stress.
A variety of modified tests can be conducted in a conventional triaxial apparatus. Bishop and Henkel (1962) proposed several modified triaxial test. To simulate field conditions, a test can be performed by keeping the axial stress constant, while decreasing the confining pressure. Consolidation can be conducted under hydrostatic condition or at any ratio of axial-to-lateral stress. A triaxial test can be conducted at any ratio of principal stresses while keeping their mean stress constant. By conducting these tests, a wide variety of stress paths can be obtained.
Historically, the most common cyclic loading technique for investigating liquefaction resistance involves the performance of the cyclic triaxial test, as a consequence of such factors as availability of equipment and relative ease of preparing undisturbed specimens. This is in spite of wide recognition of the inability of the test to accurately represent field earthquake stresses (Seed and Idriss, 1982a). Figure 2.26 and 2.27 are a schematic drawing of the cyclic triaxial test
47


apparatus and a sample recording of load, deformation, and pore pressure response, respectively.
Cyclic strength curves such as are typically generated from cyclic triaxial data are shown in
Figure 2.28. Instructions for performance of cyclic triaxial tests may be found in Engineer
Manual 1110-2-1906 (Department of the Army, 1990).
Previous studies have demonstrated that cyclic triaxial strengths (in fact, strengths determined from any unidirectional loading test) are higher than those expected to produce equivalent effects in the field (Seed, 1976). Reduction factors were developed to adjust laboratory cyclic test strengths to estimate field liquefaction resistance. The current study and additional research efforts reported in the literature indicate that estimation of field cyclic strengths from laboratory cyclic test results may not be possible by universal application of simple factors.
Figure 2.26 Schematic of cyclic triaxial test equipment (Marcuson and Krinitzsky, 1976)
48


Figure 2.27 Typical analog recordings of load, deformation, and pore pressures during a cyclic triaxial test (Department of the Army, 1990)
49


or
«✓*>
to
lii
ccr
t;
0.6
I 1 II I I 111
lOt DOUBLE AMPLITUDE STRAIN

0.4 -
100% PORE PRESSURE RESPONSE
0.2 -
SUMMARY CURVE MONTEREY SAND NO. 0 WET TAMPING COMPACTION OR * 60 PERCENT
----1---1 I I M III
1 I rTTTTT
10
J—i i i i i ii
100
NUMBER OF CYCLES
Figure 2.28 Cyclic triaxial strength curves for Monterey No.O sand (Department of the Army, 1990)
Cyclic Hollow Cylinder Tests
A hollow cylindrical test device (HCTD) is an extremely valuable tool for studying constitutive behavior under generalized stress conditions. The HCTD allows independent control of the magnitudes of the three principal stresses and rotation of the major-minor principal stress axes while recording the specimen deformational and pore pressure responses.
The University of Colorado at Denver Hollow Cylinder Torsional /Axial test cell was designed and fabricated by Dr. Jing-Wen Chen while conducting his doctoral research at UC Denver in 1988. In the hollow cylinder test at UC Denver, a hollow cylindrical soil specimen is enclosed in between an inner membrane and an outer membrane. The confining pressure can be
50


independently applied on both inner and outer chambers; therefore, inner and outer pressures can be controlled either equally or unequally. The axial load and torque are applied on the top of specimen and transmitted by a top cap or a pedestal to the specimen (Jing-Wen Chen, 1988).
When each of these boundary stresses can be controlled independently, both the principal stress direction and the relative magnitude of the intermediate principal stress can be controlled, thus the hollow cylindrical apparatus (HCA) can facilitate more generalized stress path testing than the conventional test apparatus. It is also possible to control (or measure) the pore water pressure and apply back pressure, so that drainage conditions can be controlled and both drained and undrained tests can be performed. As a result, the HCA offers an opportunity of extending the stress path approach to include simulation of both principal stress rotation and variation in intermediate principal stress, as well as conducting fundamental research into the effect of principal stress rotation under a reasonably generalized stress state.
Idealized stress conditions in a hollow cylindrical element subjected to axial load, W, torque, Mt, internal pressure, Pi, and external pressure, P0.
During shearing, the torque, Mt, develops shear stresses, Toz and xze (xoz = xze) in vertical and horizontal planes, the axial load, W, contributes to a vertical stress, oz. Pi and P0 establish a gradient of radial stress, or, across the cylinder wall. The relationship between radial stress, or, and the circumferential stress, 00, is expressed by the equilibrium equation:
00 = Or + r (dor / dr)
where r is the radial distance to a point in the hollow cylinder, and dor and doe are the radial and circumferential stress increments respectively. The stress condition in an element of a hollow cylinder specimen is shown in Fig. 2-29. Both inner and outer pressure are applied on the
51


membrane so that there is no shear stress on the vertical boundaries, or is always a principal stress because there are no shear stresses on circumferential surface throughout the wall.
Figure 2.29 Idealized stress and strain components within the HCA subjected to axial load, W, torque, Mr, internal pressure, Pi, and external pressure, Po: (a) hollow cylinder coordinates; (b) element component stresses; (c) element component strains; (d) element principal stresses (after Zdravkovic and Jardine, 2001)
The state of stress in a hollow cylinder test is defined with reference to cylindrical coordinates, in terms of the stress components.
52


°r
0
0
0 0

T-.e a.
Since the stresses will not be uniform across the wall of the cylinder for various loading conditions, to consider the hollow cylinder as an element, it becomes necessary to calculate average stresses, oz, or, 00 , toz . Hight el al. (1983) used the following expressions:
Average vertical stress oz = [W/rc (r02 - n2)] + [(P0r02 - Pin2) / (r02 - n2)]
Average radial stress or = (P0r0 + Pin) / (r0 + n)
Average circumferential stress oe = (P0r0 - Pin) / (r0 - n)
Average shear stress ioz = 3Mt / 2n (r03 - n3)
In hollow cylinder tests, the radial stress, or, is usually equal to the intermediate principal stress (02). The major and minor principal stresses, 01 and 03, are observed from the average stress components oz, 00, and ioz, and as following:
01 = [(Oz + O0 )/2] + V{[Or -O0 )/2]2 + (X0Z )2}
02 = Or
03 = [(oz + 00 )/2] -VOr -00 )/2]2 + (X0Z )2}
By regarding the specimen as a single element, the state of strain is presented in
cylindrical coordinates in terms of the following components:
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0 0
0 Yez
2
0 Yze £z
2
Also, it is necessary to calculate the average strains. According to the paper of Hight et al. (1983), the average strains are calculated using the following equations:
Average axial strain e: = AH/H
Average radial strain sr = - [(u0-ui)/(r0-ri)]
Average circumferential strain se = - [(u0+ui)/ (r0+n)]
Average shear strain y$: = [29 (r03-n3)] / [3H (r02-ri2)]
Since the average values of e- and ye: are based on strain compatibility only, the expressions for the average strains are valid and independent of the constitutive law of the material. The average values of sr and ee are based on a linear variation of radial displacement across the wall of the specimen. In the hollow cylinder test, the radial strain (sr) is usually the intermediate principal strain, sj. The major and minor principal strains can be observed from the average strain components:
£j = [(e- +£e )/2] + V{[(e- -ee )/2]2 + [yez /2]2}
52 = Sr
53 = [(or +Se )/2] -V{[(or -£e )/2]2 + \yo-_ /2]2}
Parameters a and b are two variables of stress path to describe fundamentally different aspects in the applied state of state of stress, a , is the inclination of major principal stress
54


direction with respect to the vertical axis, which can be varied from 0 to 90°. It can be computed from the known average stress components tan 2 a = 2 T0Z / (or -oe )
b is defined as the relative magnitude of the intermediate principal stress, which can be varied from 0 to 1:
b = (02-03)/(0l-03)
For the particular case of equal internal and external pressure, Pi=P0=P, and are usually assumed to be equal to P. From Average radial stress or = (P0r0 + Pin) / (r0 + n), 02 is equal to P as well. Therefore, changes in the a angle are accompanied by changes in magnitude of b. When Pi=P0
b = sin2 a (Hight. etal., 1983)
The direction of strain increment a* can be calculated from the incremental strain components tan2ade = d yo-_'/ (d sz -d ee)
The amount of non-coaxiality was defined as the difference between the directions of principal stress and of principal strain increments as, a* -a.
Cyclic Simple Shear Test
Cyclic simple shear tests were conducted on Drammen clay by Andersen et al. (1980) to investigate the effects of shear stress on clay through the overconsolidation ratio, one-way cyclic loading, shear stress amplitude which varied during the test, strain-controlled cyclic loading, horizontal shear stress during consolidation, drainage during and after undrained cyclic loading, and artificial cementation. In case of cyclic loading, the test can be used in evaluating the liquefaction potential of sands.
The cyclic simple shear test is capable of reproducing earthquake stress conditions
55


much more accurately than the cyclic triaxial test (Figure 2.30). It is more commonly used for liquefaction testing. In this test, a short cylindrical specimen is restrained against lateral expansion by rigid boundary platens, a wire reinforced membrane or with a series of stacked rings. By applying cyclic horizontal shear stresses to the top or bottom of the specimen, the test sample is deformed in the same way as an element of soil subjected to vertically propagating S waves.
Peacoch and Seed (1968) used cyclic simple shear tests to study liquefaction problems. Subsequently, Seed and Peacoch (1971), Finn (1971, 1985), Pickering (1973), Martin et al. (1975) performed considerable cyclic simple shear tests to study liquefaction problem. In recent years, simple shear devices that allow independent control of vertical and horizontal stresses have been developed.
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Shaking Table Test
The other type of laboratory test is model test. Model test usually attempt to reproduce the boundary condition and material property in the field by a small-scale physical model. It may be used to evaluate the performance of a prototype or verify predictive theories. Model test is a useful method for studying dynamic behavior of earth structure and foundation.
Shake table tests of many sizes are being used for liquefaction studies on saturated soil samples prepared in a container, fixed to a shaking platform and vibrated at the desired frequency for a prescribed time. A surcharge is placed on the sample to provide the confining pressure. The measurements of acceleration pore water pressure and settlements are made during the test. Shaking tables utilize a single horizontal translation degree of freedom, but shake table with multiple degrees of freedom have also been developed. Kokusho (1987) developed a numerical model based on shake table test.
Centrifuge Test
A geotechnical centrifuge is used to accurately conduct model tests in studying geotechnical problems such as strength, stiffness and capacity of foundations for bridges and buildings. It makes use of centrifugal acceleration to match soil stresses in a 1/50 scale model.
So, for a model container 1 m deep filled with soil, subjected to a centrifugal acceleration of 50 g, the pressures and stresses will be increased by that factor of 50. The purpose of the centrifuge machine is to shake the receptor in a controlled manner to simulate a dynamic event similar to an earthquake. But, most importantly, it is useful to study ground-shaking effects without risking public safety.
The basic principle of centrifuge modeling is to recreate the stress conditions which would exist in a full scale construction (prototype), using a model on a greatly reduced scale.
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This is done by subjecting the model components to an enhanced body force, which is provided by a centripetal acceleration of magnitude “ng”, where g is the acceleration due to the Earth Gravity (i.e. 9.81 m/s2). Stress replication in an nth scale model is achieved when the imposed “gravitational” acceleration is equal to “ng”. Thus, a centrifuge is suitable for modeling stress-dependent problems. Moreover, reduction of time for model tests such as consolidation time can be achieved by using a reduced size model.
Factor Safety against Liquefaction
The final step in the liquefaction analysis is to calculate the factor of safety against liquefaction. If the cyclic stress ratio caused by the anticipated earthquake is greater than the cyclic resistance ratio of the in situ soil, then liquefaction could occur during the earthquake, and vice versa. The factor of safety against liquefaction (FS) is defined as follows:
FS = CRR/CSR
The higher the factor of safety, the more resistant the soil is to liquefaction. However, soil that has a factor of safety slightly greater than 1.0 may still liquefy during an earthquake.
If it is determined that the soil has the ability to liquefy during an earthquake and the soil is below or will be below the groundwater table, then the liquefaction analysis is performed. The first step in the simplified procedure is to calculate the cyclic stress ratio, also commonly referred to as the seismic stress ratio (SSR) that is caused by the earthquake.
To develop the CSR earthquake equation, it is assumed that there is a level ground surface and a soil column of unit width and length, and that the soil column will move horizontally as a rigid body in response to the maximum horizontal acceleration amax exert by the earthquake at ground surface. The weight W of the soil column is equal to ytz, where yt = total
58


unit weight of the soil and z = depth below ground surface. The horizontal earthquake force F acting on the soil column (which has a unit width and length) is:
F = ma = (W/g) a = (ytz/g) amax = oV0 (amax/g)
where F= horizontal earthquake force acting on soil column that has a unit width and length, lb or kN
m = total mass of soil column, lb or kg, which is equal toW/g.
W = total weight of soil column,lb or kN. For the assumed unit width and length of soil column, the total weight of the soil column is ytz
yt = total unit weight of soil, Ib/ft3 or kN/m3 z = depth below ground surface of soil column
a = acceleration, which is the maximum horizontal acceleration at ground surface caused by the earthquake (a=amax), ft/s2 or m/s2
amax = maximum horizontal acceleration at ground surface that is induced by the earthquake, ft/s2 or m/s2.
Ovo = total vertical stress at bottom of soil column, Ib/ft2 or kPa. The total vertical stress =
ytz
Since the soil element is assumed to have a unit base width and length, the maximum shear force F is equal to the maximum shear stress Tmax,
Tmax — F — Ovo (amax/g)
dividing both sides of the equation by the vertical effective stress ovo’ gives
(Tmax / CJvo ) — (Ovo / Ovo )(amax/g)
Seed and Idriss (1971) incorporated a depth reduction factor yd, showed in Figure 2-31, into the above equation: (xmax / oV0’) = yd (oV0 / oV0’ )(amax/g)
59


For the simplified method, Seed et al. (1975) converted the typical irregular earthquake record to an equivalent series of uniform stress cycles by assuming the following:
Tcyc = 0.65t max
where xcyc = uniform cyclic shear stress amplitude of earthquake (Ib/ft2 or kPa)
CSR — (Xmax / Ovo ) — 0.65yd (Ovo / C5vo )(amax/g)
Another option is to assume a linear relationship of yd versus depth and use the following equation (Kayen et al. 1992): yd = l-0.012z
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
Stress reduction factor, rd
Figure 2.31 Reduction factor to estimate the variation of cyclic shear stress with depth below level or gently sloping ground surfaces. (After Seed and Idriss, 1971)
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CHAPTER III
LITERATURE REVIEW IUFACTOR LIQUEFACTION RESISTANCE OF SOILS
General
A monograph on the subject of ground motions and soil liquefaction by Seed and Idriss (1988a) provided a comprehensive list of the factors most often studied as influential on cyclic strength of soil, divided into three categories: (1) soil properties, including dynamic shear modulus and damping characteristics, unit weight, grain characteristics, relative density and soil structure (fabric); (2) environmental factors, such as mode of soil deposition, seismic history, geologic history (aging), coefficient of lateral earth pressure at rest, K0, over consolidation ratio, depth to water table, and effective confining pressure; and (3) earthquake characteristics, specifically ground shaking intensity and duration.
Effects of Soil Factors
Soil Properties
In term of the soil types most susceptible to liquefaction, Ishihara (1985) stated: “the hazard associated with soil liquefaction during earthquakes has been known to be encountered in deposits consisting of fine to medium sand and sands containing low plasticity fines.” Occasionally, however, cases are reported where liquefaction apparently occurred in gravelly soils.
Thus, the soil types susceptible to liquefaction are nonplastic (cohensionless) soil. An approximate listing of conhesionless soils from least to most resistant to liquefaction is clean sands, nonplastic silty sands, nonplastic silt, and gravels. There could be numerous exceptions to this sequence. For example, Ishihara (1985, 1993) described the case of tailings derived from the mining industry that were essentially composed of ground-up rocks and were classified as rock
61


flour. Ishihara (1985, 1993) stated that the rock flour in a water-saturated state did not possess significant cohesion and behaved as if it were clean sand. These tailings were shown to exhibit as low a resistance to liquefaction as clean sand.
Seed et al. (1983) stated that based on both laboratories testing and field performance, the great majority of cohesive soils will not liquefy during earthquakes. Using criteria originally stated by Seed and Idriss (1982) and subsequently confirmed by Youd and Gilstrap (1999), in order for a cohesive soil to liquefy, it must meet all the following three criteria:
(1) The soil must have less than 15 percent of the particles, based on dry weight, that are finer than 0.005 mm (i.e., percent finer at 0.005 mm < 15 percent).
(2) The soil must have a liquid limit (LL) that is less than 35 (that is, LL < 35).
(3) The water content co of the soil must be greater than 0.9 of the liquid limit [that is, co > 0.9 (LL)].
If the cohesive soil does not meet all three criteria, then it is generally considered to be not susceptible to liquefaction. Although the cohesive soil may not liquefy, there could still be a significant undrained shear strength loss due to the seismic shaking.
Polito (2001) found that soils with LL<25 and PI<7 are liquefiable, soils with 25 62


X
a)
"O
c
o
«
a.
10 20 30 40 50 60 70 80
Liquid Limit
Fig 3.1 Data presented by Wang (1979) which led to the development of the Chinese Criteria
Fig 3.2 Characteristics of Adapazari soils that were tested in J.D.Bray & R.B.Sancio (2006)’s study: (a) location on the Casagrande plasticity chart; (b) relationship between liquid limit and moisture content
I.M.Idriss and R.W.Boulanger (2004) showed that fine-grained soils transition from sandlike to clay-like behavior at plasticity indices (PI) between about 3 and 8, with the transition point appearing to be slightly lower for ML-CL soils than for ML soils. For practical purposes, it is recommended that fine grained soils be categorized as sand-like (i.e., susceptible to liquefaction) if they have a PI < 7 and clay-like (i.e., susceptible to cyclic failure, not
63


liquefaction) if they have a PI > 7. This criterion may be adjusted on a site-specific basis if justified by the results of detailed in situ and laboratory testing.
Fine Content and Plasticity Index
In Dr. Hsing-Cheng Liu’s dissertation (1992), it found that inclusion of fines in a clean sand at a constant overall void ratio does not necessarily cause increase in liquefaction resistance, and also at a constant overall void ratio, the increase in fines content at a constant plasticity index in the clean sand cause a decrease in liquefaction resistance, beyond 30%, the further increase in fines content results increase in liquefaction resistance.
Nien-Yin Chang (1990) run triaxial tests with samples tested underl5 psi confining pressure in the medium-sand with 5% fines and plasticity indices ranging from 0 to 40. The triaxial tests results showed that the trend of increasing liquefaction resistance with increasing plasticity index is more obvious, and also the effect of fine contents is very much more significant than the effect of plasticity indices.
I.M.Idriss and R.W.Boulanger (2004) showed that cyclic stress ratio (CSR) increased with increased fine content percent under the same modified standard penetration blows count. The cases for cohesionless soils with FC > 35% are plotted in Figure 3.3. Figure 3.4 shows the case history points for cohesionless soils with 5 %< FC <15%, while Figure 3.5 shows the cases for 15% < FC < 35%.
However, Chang (1990) showed the clean medium sand has the strongest liquefaction resistance and as the fine content increases, the liquefaction resistance decreases until the fine content reaches approximately 26%, as indicated by the liquefaction potential curves shifting toward the left in Figure 3.6. Then the trend reverses itself and soils begin to gain strength as the
64


fine content further increases, as indicated by the curves shifting toward the right. In Figure 3.6 shows, for the medium-sand test series, exactly the similar trend of decreasing resistance and then increasing resistance with the increasing fine content. Each curve in Figure 3.6 gives the stress ratios required to achieve initial liquefaction in ten cycles of loading for soils with the same plasticity index for each curve. The collection of curves also indicates the existence of a critical fine content, below which, the resistance decrease, and, above which, the resistance increases with increasing fine content, and the critical fine content decrease with increasing plasticity of soils. Figure 3.6, 3.7, and 3.8 seem to indicate that the critical fine content increases as the number of cycles required reaching initial liquefaction increases.
Tzuo-Shin Ueng, Chia-Wen Sun and Chieh-When Chen (2013) showed that as the fine content increases, cyclic resistance ratio decreases until the fine content approximately 20% in the Figure 3.9. And also indicated that the effect of fines on the liquefaction resistance of a soil is more prominent using (FC)4oo (passing the No.400 sieve) than (FC)2oo (passing the No. 200 sieve), probably due partly to the different in plasticity of the fines.
Yong Wang and Yanli Wang (2010) showed that with a constant dry density, the liquefaction resistance first decrease and then increase with the increase of the fines content, when the relative density reaches the minimum value at the fines content of 30% and the liquefaction resistance also reaches the minimum value at the same fines content in Figure3.10.
Mehmet Murat Monkul and Jerry A. Yamamuro (2011) showed that relative density alone cannot be a consistent comparison basis for the influence of fines content on liquefaction potential of sand in Figure 3.11, and also the Dso-sand/dso-siit ratio becomes larger (for SilCoSil and Potsdam fines), the void ratio decreases consistently with increasing fines content, show as Figure 3.12.
65


Chang (1990) showed that the medium sand with a small fine content of 5% reflects in the irregular relationship between the stress ratio required to reach initial liquefaction versus plasticity index, and also comparison between the effect of fine contents and the effect of plasticity indices of fines on the liquefaction resistance of soils indicated that the effect of fine contents is very much more significant than the effect of plasticity index.
Modified Standard Penetration • (N^qq - Blows/ft
Fig. 3.3 SPT case histories of cohesionless soils with FC > 35% and the NCEER Workshop (1997) curve and the recommended curves for both clean sand and for FC = 35% for M = I'A and o Vo = 1 atm (I.M.Idriss and R.W.Boulanger, 2004)
66


Modified Standard Penetration - (A/r)g0 - Blows/ft
Fig. 3.4 SPT case histories of cohesionless soils with 5% 67


Modified Standard Penetration - - Blows/ft
Fig.3.5 SPT case histories of cohesionless soils with 15% 68


Figure 3.6 Stress Ratio Required to reach Liquefaction in 10 Cycles versus Fine Content Under Confining pressure 15 psi (N.Y. Chang, 1990)
Figure 3.7 Stress Ratio Required to reach Liquefaction in 30 Cycles versus Fine Content Under Confining pressure 15 psi (N.Y.Chang, 1990)
69


o
8
O
0.00 8.00 16.00 2U.00 32.00 U0.00 U8.00 56.00 6*1.00
FINE CONTENT (7.)
Figure 3.8 Stress Ratio Required to reach Liquefaction in 100 Cycles versus Fine Content Under Confining pressure 15 psi (N.Y.Chang, 1990)
Figure 3.9 Cyclic resistance ratio and Chieh-When Chen, 2013)
versus equivalent (FC)4oo (Tzuo-Shin Ueng, Chia-Wen Sun
70


Fig.3.10 Variation in Liquefaction Resistance with Fines Content for Ni=20 (Yong Wang and Yanli Wang, 2010)
FC (%)
Fig. 3.11 Change of relative density and liquefaction potential with different fines contents and Silts for tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011)
71


Fig.3.12 Change of void ratio and liquefaction potential with different fines contents and silts for tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011)
Soil Relative Density Dr
Based on field studies, cohensionless soils in a loose relative density state are susceptible to liquefaction. Loose nonplastic soils will contract during the seismic shaking which will cause the development of excess pore water pressure.
For dense sands, the state of initial liquefaction does not produce large deformations because of the dilation tendency of the sand upon reversal of the cyclic shear stress. Poulos et al. (1985) stated that if the in situ soil can be shown to be dilative, then it need not evaluated because it will not be susceptible to liquefaction.
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LIQUEFACTION RESISTANCE OF MONTEREY NO.0/30 SAND CONTAINING FINES UNDER CYCLIC TRIAXIAL AND CYCLIC HOLLOW CYLINDER TESTS by JUNGANG LIU B.S., Wuhan University of Technology, China, 2007 M.S ., University of Colorado Denver , 2012 A thesis submitted to the Faculty of the Graduate School of the University of Colorado Denver in partial fulfillment of the requirements for the degree of Doctor of Philosophy Civil Engineering Program 2019

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© 2019 JUNGANG LIU ALL RIGHTS RESERVED

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This thesis for the Doctor of Philosophy degree b y Jungang Liu Has been approved for the Civil Engineering Program by Shideh Dashti, Chair Nien Yin Chang, Advisor Dobroslav Znidarcic Brian Brady Nghiem Hien Shingchun Wang Date: May, 18 2019

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Liu, Jungang (Ph. D, Civil Engineering) Liquefaction Resistance of Monterey No. 0/30 Sand Containing Fines under Cyclic Triaxial and Cyclic Hollow Cylinder Tests Thesis directed by Professor Nien Yin Chang ABSTRACT Liquef action is the most detrimental ground failure caused by strong earthquakes. Ground liquefaction leads to associated foundation and superstructure failures due to loss of bearing capacity and excessive deformation and an appropriate assessment of liquefaction is critical to the seismic safety evaluation of foundations and super structures. This study concerns the liquefaction resistance of Monterey No. 0/30 sand contacting fines. Monterey No. 0/30 sand is a uniform clean medium sand. Fines were prepared by sieving Leyden clay through a U.S. # 200 sieve. One hundred fourteen isotopically consolidated undrain cyclic triaxial tests and thirty seven cyclic hollow cylinder test were performed to investigate the effect of fines content on liquefaction resistance of soils. This research includes compare and relate the soil liquefaction resistance found by cyclic triaixal and cyclic hollow cylinder test results on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content. The cyclic triax ial and cyclic hollow cylinder liquefaction resistance s of soils were expressed in terms of liquefaction potential curves. From the liquefaction potential curves, stress ratio in cyclic triaxial and cyclic hollow cylinder tests causing initial liquefaction in 10 cycles, 30 cycles, 40 cycles and 50 cycles were chosen as dependent variables in the regression model. Three independent variables of regression models, deviator stress in cyclic triaxial test, DS (cyclic shear stress in cyclic hollow cylinder test, CSS); fine content (decimal), FC; consolidation pressure, CP , were eventually selected for final statistical analysis.

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In addition to the evaluation of liquefaction potential, an excess pore pressure generation was simulated using Horita constitutive mod el with the parameters evaluated by laboratory tests . This research includes c omparison of excess pore pressure generation between simulations from can be ap plied to better understand the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils. The form and content of this abstract are approved. I recommend its publication. Approved: Nien Yin Chang

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ACKNOWLEDGEMENTS I would like to express my sincerest appreciation to my advisor, Professor Nien Yin Chang for his constant support and guidance throughout the course of this study. Gratitude is extended to Dr. Shideh Dashti, Dr. Dobroslav Znidarcic, Dr. Brian Brady, Dr. Nghiem Hien and Dr. Shingchun Wang for serving on my final examination committee . I would like to thank my student colleague, Mr. Brian Volmer for his friendship and help in preparing this study. Thanks are extended to Mr. Tom Thuis and his staff in Calibration shop at the University of Colorado Denver for their assistance in instrume ntation. I would like to thank my wife, Liang Feng for many years of her endurance, sacrifice, assistance and encouragement, which made this accomplishment possible. I also thank my parents for their understanding and support.

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TABLE OF CON TENTS CHAPTER 1. 2. LITERATURE REVIEW: 1995 Hyogoken 1999 Chi Gene Othe r Techniques ( Laboratory Methods for Soil Liquefaction Undisturbed Sam

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Laboratory T Cyclic Triaxi al Cyclic Hollow C Shaking Table T Centrifuge Factor Safety against Liq 3. LITERATURE REVIEW: FACTORS AFFECTING LIQUEFACTION RESISTANCE Effects of Soil F ac Soil Proper Fines Contents and Plasti Soil Relative D Particle Size Gradation Particle Sh Geological Aging and Ceme Effects of Laboratory Specimen Preparatio Reconstitution versu s Int Loading Wave Frequency on Cycli Specimen

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Frictionless Caps an Relative Den 3 Cyclic Stress Amplitude and Number of C Particles Size and Grad Pre Lateral Earth Pressure (K 0 ) and Over Consolidation Ratio, (K c Effects of Field Fact Groundwater Placement Conditions or Deposi Drainage Condit Confining Pre Historical Envir Building Effects of Earthquake Magnitude Sc Intensity and D Ground Motion 4. CYCLIC TRIAXIAL TEST PROGRAM AND Introduction Test Program

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Test Equipmen Soil Samples Prepara Soil Samp Mixing Soil S Samples Prepar Samples Quality C Relative Densit Determining B Param Test Procedure Tests Results Soil Liquefaction Potent Cyclic Axial load versus Number of Cycles to Excess Pore Water Pressure versus Number of Cyclic Deviator Stress versus Effect of Fines Content o 137 5. HOLLOW CYLINDER TEST APPARATUS, PROGRAMS AND TEST RESULTS Hollow Cylinder Test Ap Principles of Hollow Cylinder Testing Stress Distribution in Hollow C

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Specimen Geo Membrane Penetratio Soil Samples Prepara Soil Samp Mixing Soil S Samples Prepar Samples Quality C Relative Densit Determining B Test Procedure Analysis of Results. ......................................................................................................... 176 Cyclic Torque versus Number of Cyc Excess Pore Water Pressure versus Number of Cyclic Shear Stress versus S Stress Pa Effect of Fines Content on Lique 6. COMPARISON OF CYCLIC TRIAXIAL AND HOLLOW CYLINDER TEST Correction Factor between Cyclic Stress Ratio Causing Liquefaction in the

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Field and Cyclic Stress Ra tio Causing Liquefaction of Triaxial Test Sample in the .185 Comparison on both test re Cyclic Torque, Cyclic Axial Load versus Numb Pore Water Pressure versus Number of C Cyclic Shear Stress, Cyclic Deviator Stress versus Shear Strain, Axial Strain...196 7. Previous Th r eshold Fines Conten Definitio Factors Effects Threshold Fines C Results of laboratory Tes All Soil Samples Results in Cyclic Triaxial All Soil Samples Results in Cyclic H 8. EXCESS PORE PRESSURE Excess pore pressure generation from lab Excess pore pressure Generation from la Constitutive models for simulating pore pr Constitute Model UBC3D Finn Constitutive model

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9. STATISTICAL MODELING OF LIQU Selection of Variabl Initial Variable Se Variable Reduc Final Variable Sele Regression Model for Liquefaction Regression model for cyclic triaxial test Regression model for cyclic hollow 10. A PROPOSED NEW PROCEDURE FOR EVALUATING LIQUEFACTION RESI STANCE OF SOIL WITH P Procedure by Seed e Proposed Procedu re Cyclic Resistance Ratio from S SPT based Case History Cases History from Kohji Tokimatsu and Yosh iaki Yoshimi.................278 Proposed Procedure Cyclic Resistance Ratio from SPT bas ed laboratory testing data...286 Calculating SPT Blow Count (N 1 ) 60 Calculation of SPT Blow Count based Laboratory Test Data........... ......289 Comment on the New Pr 11. SUMMARY CONCLUSION AND RECOMMENDATIONS FOR FUTURE

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Sum Recommendations for Future A. Field Methods for Soil Liquefaction Resis B. Factors Affecting Liquefaction Resis C. Procedure for Fabricating Rubber D. Cyclic Triaxial and Hollow Cylinder T

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LIST OF TABLES TABLE 1. Summary of CPT Field case histories from the 1995 Hyogoken Nambu (Kobe) earthquake. 2. 3. 4. Relationship between (N 1 ) 60 and po 5. 6. Estimated susceptibility deposits to liquefaction during strong seismic shaking based on geologic age and depositional environment. (M 7. All soil samples in cyclic triaxial test (Jun gang Liu, 2019 8. Test program details in hollow cylinder test (Jungang Liu, 2019 9. Result of comparati ve study on clean sand ( Liu, 10. 11. Values of C r (some resear 12. Cr values on the clean sand produced in Labo ratory by CTT and CHCT ( Liu, et al, 201 7 13. C r values on the mixture samples produced in laboratory by CTT and CHCT (Jungang Liu, 2019 14. ...189 15. Threshold fines content in cyclic triaxial test and cyclic hollow cylinder test (Jungang Liu, 2019

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16. Correlation coefficient matrix of threshold fines content and laboratory factors (Jungang Liu, 2019 17. All parameters in Horita model for trail 1 (Masakun 18. All parameters in Horita model for trail 2 (Masakun 19. Correlation coefficient matrix of the dependent variable , SRCTT10 and the sixteen chosen primary independent variables (Jungang Liu, 2019 20. Correlation coefficient matrix of the dependent variable, SRCTT30 and the sixteen chosen primary independent variables (Jungang Liu, 2019 256 21. Correlation coefficient matrix of the dependent variable, SRCTT40 and the sixteen chosen primary independent variables (Jungang Liu, 2019 22. Correlation coefficient matrix of the dependent variable, SRCTT50 and the sixteen chosen primary i ndependent variables (Jungang Liu, 2019 23. Correlation coefficient matrix of the dependent variable, SRCTT30 (Dr30) and the sixteen chosen primary independent variables (Jungang Liu, 2019 24. Correlation coefficient matrix of the dependent variable, SRCTT30 (Dr45) and the sixteen chosen primary independent variables (Jungang Liu, 2019 25. Correlation coefficient matrix of the dependent variable, SRCHCT10 and the sixteen chosen primary independent variables (Jungang Liu, 2019 26. Correlation coefficient matrix of the dependent variable, SRCHCT30 and the sixteen chosen primary independent variables (Jungang Liu, 2019 27. Summary of SPT based liquefaction case history data from Idriss an d Boulanger (2004,

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28. Case histories data from Idriss and Boulanger: (N 1 ) 60 and CSR with different Fines Content data were created by Jungang Liu ( 2019 29. Summary of SPT based liquefaction case history data from Tokimatsu and Yoshimi 30. Case histories data from Idriss and Boulanger: N 1 and CSR with different Fines Content data were created by Jungang Liu ( 2019 31. Summary of Rod Ene rgy Ratios (Skemp

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LIST OF FIGURES FIGURE 2.1 Niigata city destroyed by the Niigata Earthquake 1964. (National Geophysical Data Center).5 2.2 Soil profile at Kawagishi cho 2.3 Cracked Highway in 2.5 Location 2.6 Map Showing CPT Field Case History Locations from the 1995 Hyogoken Nambu Earthquake, in relation to the SPT Case History Locations and PGA Contours from Cetin et 2.7 Proposed CPT based liquefaction boundary curve (Ku, Chih Sheng, Der Her Lee, and Jian 2.8 Observed liquefaction in the park (left) and parking area of Disneyland (right) . (S. Bhattacharya, et, 20 2.10. Ground Failures and Resulting Water Damage. (Christchurch, New Zealand Earthquake 2.11 Liquefaction documentation map of eastern Christchurch from drive through Reconnaissance. (C 2.12 Relationship between cyclic stress ratios causing liquefaction and (N 1 ) 60 values for clean Sands in M 2.13 Re lationship between cyclic stress ratios causing liquefaction and (N 1 ) 60 values for silty Sands in M 2.14 Updated SPT case history database of liquefaction in cohesionless soils with v arious fines v = 1 atm and equivalent clean sand (N 1 ) 60cs (Idriss 2.15 SPT case history database used previously by Idriss and Boulanger (2 2.16 Parts (a) and (b) of Figure 2.6 in Cetin et al. (2004); note that the points representing case Histories are identical in parts (a) and (b) of this figure and as representing conditions with

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2.17 Curves relating CSR to (N 1 ) 60 published over the past 24 years for clean sands and the Recommended curve for M ' vo 2.18 Manufacturing and operating 2.19 Schematic section through a piezocone head, showing the piezo element and friction sleeve 2.20 CPT prediction of overburden pressure corrected SPT blow count (Olsen and Malone, 2.21 CPT based case histories and recommended relation for clean sands for M ' vo = 1 atm (I.M. Idriss and R.W. Boula 2.22 CPT based case histories and recommended relation for clean sands with relations 2.24 2.25 Results of the SCPT carried out on June 16th 2001 in a permafrost mound near Umiujaq, 2.26 Sche matic of cyclic triaxial test equipment (Marcus 2.27 Typical analog recordings of load, deformation, and pore pressures during a cyclic triaxial Test (Department of the Ar 2.28 Cycl ic triaxial strength curves for Monterey No.0 sand (De partment of the Army, 1990)...50 2.29 Idealized stress and strain components within the HCA subjected to axial load, W , torque, M T , internal pressure, Pi , and external pressure, Po : (a) hollow cylinder coordinates; (b) Element component stresses; (c) element component strains; (d) element principal stresses (After Zdravkovic and Jardi 2.30 Cyclic 2.31 Reduction factor to estimate the variation of cyclic shear stress with depth below level or Gently sloping ground surfaces. (After Seed 3.1 Data presented by Wang (1979) which led to the developm 3.2 (a) Location on the Casagrande plasticity chart; (b) relationship between liquid limit and

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3.3 SPT case histories of cohesionless soils with FC Curve and the recommended curves for both clean sand and for FC = 35% for M ' vo = 1 atm (I.M.Idriss and R.W.Boula 3.4 SPT case histories of cohesionless soils with 5%
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3.18 Amplification attenuation relationship for modifying bedrock acceleration at soft soil site s 4.1 Test Specimen Placed Within the Triaxial Cell. (Jungang Liu, 2019 4.2 Pore pressure transducer and a 4 way valve used to monitor cell pressure, back pressure, and Speci men pore water pressure. (Jungang Liu, 2019 4.3 Pressure control panel used to apply cell and back pressure to the triaxial cell. Graduated Burettes were used to determine specimen volume change. (Jungang Liu, 2019 1 4.4 Closed loop electro hydraulic materials test system applying a sinusoidal loading to Soil specimen. (Jungang Liu, 2019 2 4.5 Rubber o ring (Jungang Liu, 2019 4.6 Vernier caliper used to measure the thickness of the membrane and diameter of the sand Specimen (Jungang Liu, 2019 4.7 Vernier caliper used to measure the height of the test specimen. (Jungang Liu, 2019 6 4.8 Looking down on the base of the triaxial cell. (Jungang Liu, 2019 4.9 Rubber membrane placed on the base of the triaxial cell. (Jungang Liu, 2019 4.10 Cylindrical mold placed on the base of the triaxial cell. (Jungang Liu, 2019 8 4.11 The membrane adhereon the inner wall of the mold. (Jungang Liu, 2019 8 4.12 Soil sample placed within the zero raining device. (Jungang Liu, 2019 119 4.13 Porous stone placed on top of the Monterey No. 0 Sand Specimen. (Jungang Liu, 2019 ).119 4.14 Top cap placed on top of the specimen (Jungang Liu, 2019 4.15 Positioning of the top surface of the loading cap parallel to the ba se of the triaxial cell (Jungang Liu, 2019 4.16 Measuring the length of the sample using a vernier caliper (Jungang Liu, 2019 4.17 Pi tape (Jungang Liu, 2019 4.18 Specimen was saturating (Jungang Liu, 2019 4.19 Set up on the computer to run MTS. (Jungang Liu, 2019

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4.20 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 30% and effective stress 15psi. (Jungang Liu, 2019 4.21 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different perc ent of fines contents at relative density of 30% and effective stress 30psi. (Jungang Liu, 2019 4.22 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 45% and effective stress 15psi. (Jungang Liu, 2019 4.23 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 45% and effective stress 30psi. (Jungang Liu, 2019 4.24 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 60% and effective stress 15psi. (Jungang Liu, 201 9 4.25 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with Different percent of fines contents at relative density of 60% and effective stress 30psi. (Jungang Liu, 20 19 3 4.26 Cyclic Axial Load versus Number of cycles to liquefaction in Cyclic Triaxial Test. (Jungang Liu, 2019 134 4.27 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. (Jungang Liu, 2019 4.28 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. (Jungang Liu, 2019 4.29 P versus q stress path in Cyclic Traixial Test. (Jungang Liu, 2 019 5.1 Idealized stress and strain components within the HCA subjected to axial load, W , torque, M T , internal pressure, Pi , and external pressure, Po : (a) hollow cylinder coordinates; (b) Element component stresses; (c) element component strains; (d) element principal stresses (After Zdravkovic and Jardin 5.2 Definitions of average stresses and strains (after Hight et al., 5.3 Definitions used for stress non uniformity and accuracy (after Hight et al., 5.4 Effect of stress ratio level on non uniformity coefficients (after Vaid et al. 5.5 Shear stress distribution in hollow cylinder torsional sh ear test specimens (after Porovic,

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ss in HCTA s in HCTA 5.8 Schematic diagram of the HCTA 88 at the University of Colorado at Denver Geotechnical Laboratory (Chen, 19 5.9 Bottom Platen (Jungang Liu, 2019 5.10 Base Plate (Jungang Liu, 2019 5.11 Bottom Sample Pedestal (left) and Top Sample Pedestal (Jungang Liu, 2019 5.12 Top Cap and Top Plate (Jungang Liu, 2019 5.13 Piston Locking Mechanism (Jungang Liu, 2019 5.14 Supporting Bars (Jungang Liu, 2019 5.15 Pressure Chamber (Jungang Liu, 2019 5.16: Hollow cylinder test setup in software of MTS. (Jungang Liu, 2019 5.17 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test. (Jungang Liu, 2019 5.18 Pore water pressure change versus Number of cycles to liquefaction in CHCT. (Jungang Liu, 2019 5.19 Cyclic shear stress versus shear strain in Hollow Cylinder Test. (Jungang Liu, 2019 5.20 p versus q stress path in Hollow Cylinder Test. (Jungang Liu, 2019 6.1 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test ( Liu, et al, 2017 .......191 6.2 Cyclic Axial Load versus Number of cycles to liquefaction in Cyclic Triaxial Test. ( Liu, et al, 2017 ..192 6.3 Pore water pressure change versus Number of cycles to liquefaction in HC Test. ( Liu, et al 2017 194 6.4 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. ( Liu, et al, 2017 .....195

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6.5 Cyclic shear stress versus shear strain in Hollow Cylinder Test ( Liu, et al, 2017 6.6 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. ( Liu, et al, 2017 6.7 p versus q stress path in Hollow Cylinder Test. ( Liu, e t al, 2017 6.8 p versus q stress path in Cyclic Traixial Test. ( Liu, et al, 2017 7.1 SPT case histories of cohesionless soils with FC 35% and the NCEER Workshop (1997) Curve and the recommended curves for both clean sand and for FC = 35% for M = 7½ and ' vo = 1 atm ( I.M.Idriss and R.W.Boul 7.2 SPT case histories of cohesionless soils with 5%
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Voids and solids (b) variation on e min . (Misko Cubrinovsk 7.13 Variation in e max and e min with fine content of mixture s of Cambria Sand and Nevada fines. (Misko Cubrinovski and Kenji Ishih 7.14 Intergranular soil mix classification. (S. Thevanayagam; T. Shenthan; S. Mohan and J. 216 7.15 Stress ratio require for liquefaction in 10 and 35 cycles versus variable fine content (M L Mixing samples at relative density 30% and confining pressure 15,30 psi) in cyclic triaxial Tests. (Jungang Liu, 2019 7.16 Stress ratio require for liquefaction in 35 and 50 cycles versus variable fine content (M L Mixing samples at relative density 45% and confining pressure 15,30 psi) in cyclic triaxial Tests. (Jungang Liu, 2019 1 9 7.17 Stress ratio require for liquefaction in 50 and 90 cycles versus variable fine content (M L Mixing samples at relative density 60% and confining pressure 15,30 psi) in cyclic t riaxial Tests. (Jungang Liu, 2019 7.18 Stress ratio require for liquefaction in 8 cycles versus variable fine content (mixing samples At relative density 30% and confining pressure 15psi) in cyclic hollow c ylinder tests. (Jungang Liu, 2019 7.19 Stress ratio require for liquefaction in 27 cycles versus variable fine content (mixing Samples at relative density 30% and confining pressure 30 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019 7.20 Stress ratio require for liquefaction in 20 cycles versus variable fine content (mixing Samples at relative density 60 % and confining pressure 15 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019 223 7.21 Stress ratio require for liquefaction in 40 cycles versus variable fine content (mixing Samples at relative density 60 % and confining pressure 30 psi) in cyclic hollow cylinder Tests. (Jungang Liu, 2019 ...224 8.1. Correlation to estimate parameter F in Vucetic Dobry PWP generation model. (Vucetic and 8.2. Excess Pore Water Pressure ratio vs. Volumetric Strain (Chung, R.M., Yokel, F.Y. and 0 8.4 The normalized exces s pore water pressure ratio versus the normalized number of cyclic Stress cycle with different percent of fines content and relative densities in cyclic triaxial test.

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a) relative density at 30% and different of fines content (0%, 5%,10%,15%,25%,35% and 45%). (Jungang Liu, 2019 1 8.5 The normalized excess pore water pressure ratio versus the normalized number of cyclic Stress cycle with different percent of fines content and relative densities in cyclic hollow Cylinder test. a) relative density at 30% and different of fines content (0%, 5%,10%,15%,25% and 35%).(Jungang Liu, 2019 3 8.6 Average excess pore pressure versus normalized number of cyclic stress cycles in all Samples a) in cyclic triaxial test with different of fines content (0%, 5%, 10%, 15%, 25%, 35% and 45%).(Jungang Liu, 2019 5 8.7 Curved Equi 4 8.8 Flow Chart for Simulation on Strain Controlled Undrained Cyclic Traixial Tests (Horita, 5 8.9 Flow Chart for Simulation on Stress Contro lled Undrained Cyclic Triaxial Tests (Horita, 246 8.10 1 Comparison between Measured and Simulated Responses of Soil Samples in Generation of Pore Water Pressure. (Jungang Liu, 2019 8.10 1 Comparison between Measured and Simulated Responses of Soil Samples in Generation of Pore Water Pressure. (Jungang Liu, 8.11 Comparison between Measured and Simulated Responses of Saturated Monterey No. 0/30 Sand to Stress Controlled Undrained Cyclic Loading with stress ratio = 0.4, No. cycles to Liquefaction =10 a) Effective Stress Path (Jungang Liu, 2019 10.1 SPT case histories of cohesionless soils with FC 35% and the NCEER Workshop (1997) Curve and the recommended curves for both clean sand and for FC = 35% for M = 7½ and ' vo = 1 atm (I.M.Idriss and R.W.Boulanger, 10.2 Shear stress reduction factor, rd, relationship (I.M.Idri 10.3. Overburden correction factor (CN) relationship for CPT and SPT penetration resistances: v ' /P a = 0 v ' /P a = 0 2 along with Liao and Whitman's (1986) 10.4. Case Histories from Idriss and Boulanger (all data for soil liquefy): (N 1 ) 60 versus FC (%) Was created by Jungang Liu ( 2019 10.5. Case Histories from Idriss & Boulanger: Average (N1)60 versus FC (0% 25%) was created By Jungang Liu ( 2019

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10.6. Case Histories from Idriss & Boulanger: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average (N1)60 with different FC (0% 20%) was created by Jungang Liu ( 2019 10.7 Case Histories from Tokimatsu &Yoshimi (all data for soil liquefy): N1 versus FC (%) was Created by Jungang Liu ( 2019 10.8 Case Histories from Tokimatsu &Yos himi: Average (N1), average MM (N1) versus FC (0% 20%) was created by Jungang Liu ( 2019 10.9. Case Histories from Tokimatsu &Yoshimi: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average N 1 with different FC (0% 20%) was created by Jungang Liu ( 2019 1 10.10.SPT blow count with different fines content (0% 25%) in case histories data from Idriss & Boulanger and Tokimatsu &Yoshimi (all data for soil liquefy) created by Jungang Liu ( 2019 282 10.11 Relationship between void ratio range and fin es content of sandy soils (Misko. Cubrinovshi & Kenji. Ishiha 10.11 Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 30%, 45% and 60%. (Jungang Liu, 2019 10.12. Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 30% (Jungang Liu, 2019 10.13 Cyclic Triaxial test results for calculating (N1)60 From FC=5% 35%, prepared at Dr= 45% (Jungang Liu, 2019 4 10.14 Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 60% (Jungang Liu, 2019 5 10.15 Calculating (N 1 ) 60 versus cyclic stress ratio from cyclic traixial test on relative density of 30% under consolidation pressure 15psi with di fferent fines content. (Jungang Liu, 2019 6 10.16 Calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 30% under consolidation pressure of 30psi with different fines content (Jungang Liu, 2019 7 10.17 Calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 45% under consolidation pressure of 15 psi with different fines content (Jungang Liu, 2019

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10.18 Calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 45% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019 10.19 Calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 60% under consolidation pressure of 15 psi with different fines content (Jungang Liu, 2019 10.20 Calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 60% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019 10.21 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC =0% 35%, prepared at Dr= 30% (Jungang Liu, 2019 10.22 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC = 0% 35%, prepared At Dr= 60% (Jungang Liu, 2019 10.23 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35%, prepared At Dr= 30% and 60% (Jungang Liu, 2019 10.24 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (N1)60 Fr om FC=0% 35%, prepared at Dr= 30% (Jungang Liu, 2019 10.25 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 60% (Jungang Liu, 2019 10.26 Calculating (N 1 ) 60 from cyclic triaxial and cyclic hollow cylinder tests on relative density 30% & 60% and consolidation pressure 15psi & 30psi, (N 1 ) 60 from case histories versus Different fines content (Jungang Liu, 2019 0 9

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98 CHAPTER IV CYCLIC TRIAXIAL TEST PROGRAM and RESULTS ANALYSIS Introduction In a cyclic triaxial test, a cylindrical specimen of soil encased in a rubber membrane is placed in a chamber, subjected to a confining fluid pressure, and then applying dynamic loading axially until soil failure. The axial load may be applied to the sample through a rigid top platen. The axial force can be compression or extension: thus, the axial stress can be either major or minor princ ipal stress. Usually the top platen is laid over a porous stone that allows fluid to flow in and out of the specimen. The axial deformation of the specimen is directly monitored by the movement of the piston that is in contact with or connected to the top platen. The lateral deformation is not usually measured. Transducers are used for pore pressure measurement. In a cyclic triaxial test, a sample is consolidated under an initial isotropic confining pressure. The effective stress is kept constant and axial load is either increased (compression test) or decreased (extension test) during a test. Thus, two of three principal stresses are always equal during a test. In a compression test, the intermediate principal stress is equal to the minor principal stress; and the axial stress is equal to the major principal stress. In an extension test, the major and the intermediate principal stress are equal, while the axial stress is equal to minor principal stress. A variety of modified tests can be conducted in a con ventional triaxial apparatus. Bishop and Henkel (1962) proposed several modified triaxial tests. To simulate field conditions, a test can be performed by keeping the axial stress constant, while decreasing the effective stress . Consolidation can be conduct ed under hydrostatic condition or at any ratio of axial to lateral

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99 stress. A triaxial test can be conducted at any ratio of principal stresses while keeping their mean stress constant. By conducting these tests, a wide variety of stress paths can be obtain ed. Test Program Triaxial Testing System The triaxial cell was shown in Fig.4.1 for performing all cyclic triaxial tests. A pore pressure transducer attached to a four way valve plugged into the right side of the triaxial cell base was used to measure th e cell pressure, back pressure, and pore water pressure at the top of the specimen. The pore pressure transducer was attached to a four way valve such that by changing the position of the valve the transducer could be subjected to either the cell pressure, back pressure, or pore water pressure at the top of the specimen. Fig.4.2 shows the transducer and four way valve disconnected from the triaxial cell. On the right side of Fig.4.1, four pressure lines are seen. The upper horizontal line entering the base of the triaxial cell applied the consolidation cell pressure, with the line passing behind the cell and over to the four way valve used for transmitting the cell pressure to the pore pressure transducer. The middle line entering the base of the triaxial c ell applied the back pressure to the specimen with the lower line used for transmitting the pore pressure from the top of the specimen to the four way valve and pressure transducer. The pressure control panel shown in Fig.4.3 could be used to apply cell and back pressures simultaneously to three triaxial cells. An air compressor connected to the pressure control panel furnished a maximum air pressure of 200 psi for the pressure contro l panel. When the pressure in the air tank dropped below a specified minimum valve, a pressure switch would turn on the compressor.

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100 The graduated burettes (vertical stand pips) shown in Fig.4.3 were used to measure specimen volume changes within an accura cy of 0.1 cubic centimeter. Specimen volume change measurements were made after consolidation of the test specimen and after attaching the loading ram to the specimen top cap. Figure 4.1 Test Specimen Placed Within the T riaxial Cell. (Jungang Liu, 2019 )

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101 Figure 4.2 Pore pressure transducer and a 4 way valve used to monitor cell pressure, back pressure, and specimen pore water pressure. (Jungang Liu, 201 9 ) Figure 4.3 Pressure control panel used to apply cell and back pressure to the triaxial cell. Gra duated burettes were used to determine specimen v olume change. (Jungang Liu, 2019 )

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102 Test Equipment Basic Principles of the MST Closed Loop System A series 810, Material Test System furnished by the MST Systems Corporation, Minneapolis, Minnesota as shown in Fig.4.4 was used to apply the cyclic dynamic loading to the sand specimens. Figure 4.4 Closed loop electro hydraulic materials test system applying a sinusoidal loading to S oil specimen. (Jungang Liu, 2019 ) Hydraulic Power Supply A fixed volume pump supply fluid pressure to the system. The hydraulic power supply may be operated locally, through use of its own controls, or via the MTS remote control panel that was the case during tests that I performed.

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103 Two levels of operation are provided: an output pressure of 300 psi for the low or bypass condition and an output pressure of 3000psi for the high condition. A safety pressure control valve protects the power supply from the buildup of excessive pressure. A fluid to water heat exchanger is used by the h ydraulic power supply to maintain the reservoir hydraulic pressure below a maximum safe temperature. A temperature sensitive switch mounted on the reservoir will open and turn off the hydraulic power supply if the hydraulic fluid temperature exceeds a pred etermined limit. Hydraulic Actuator The hydraulic actuator is the force generating and/or positioning device in the system. Movement of loading piston is the direct result of the application of fluid pressure to one side of the piston. A load applied to s ome external reaction point by the piston is equal to the effective piston area times the activating pressure. Servo valve The hydraulic actuator is controlled by the opening and closing of the servo valve in response to a control signal from the valve dr iver or controller. The servo valve can open in either of two positions, thereby permitting high pressure fluid to enter into either side of the piston. This alternating application of hydraulic pressure to either side of the piston makes it possible to ap ply smooth cyclic tension and compressive loads to a test specimen. When the servo valve is opened to allow fluid to flow into one end of the cylinder, the valve on the opposite end of the cylinder is opened to provide a path for fluid to flow back to the hydraulic power supply.

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104 The rate of fluid flow through the servo valve is in direct proportion to the magnitude of the control signal. The polarity of the control signal determines which end of the actuator cylinder will receive additional fluid thereby d etermining the direction of the piston stroke. Transducers Transducers on the MTS machine sense some quantity generated by the hydraulic actuator, such as vertical load or linear displacement, and provide an output voltage directly proportional to the measured quantity. The load cell is a force measuring transduc er that provides an output voltage directly proportional to the applied load. Compressive and tensile forces are distinguished by the polarity of the output voltage. The linear displacement of the loading ram is measured by a linear variable differential transformer (LVDT). The LVDT requires A C excitation and provides an A C output. The amplitude of the output varies in direct proportion to the amount of displacement of the LVDT core. Transducer Conditioners Transducer conditioners supply excitation vol tages to their respective transducers and control the transducer output voltages to d c levels suitable for use in the control portion of the system. Output of each transducer conditioner is 10 volts, positive or negative when the mechanical input to the t ransducer equals plus or minus 100% of the selected operating range. For example if the MTS machine was set on load control and 100% of the operating range, the output voltage would be ± 10 volts when a load of ±20,000 pounds was applied. Corresponding if the MTS was set on strain control and 100% of the operating range, the output voltage would be±10 volts when a displacement of ±5 inches occurred.

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105 Soil Sample Preparation Soil Samples The soil sample is the mixture of a uniform medium Monterey No. 0/30 sand and Leyden clay (M L) from Golden Colorado. Monterey No. 0/30 sand sample is a uniform clean sand from Monterey, California with a coefficient of uniformity, C u , equals 1.6, a coefficient of curvature, C v , 1.00, and a mean grain size, D 50 , 0.45mm. The sand is classified as SP via the Unified Soil Classification System. The sand has maximum unit weight of 105.8 pcf (Ib/ft 3 ), minimum unit weight 91.7 pcf (Ib/ft 3 ), its friction angle 37 o and specific gravity 2.65. Leyden clay was sieved through a #200 si eve to remove any impurities. The specific gravity of clay is 2.67, Liquid limit of 42%, plastic limit of 22% and plastic index of 20%. The maximum dry unit weight of clay is 109 pcf (Ib/ft 3 ) with the optimum of 17%water content in the standard proctor com paction test . Mixing Soil Samples It is extremely time consuming to achieve a uniform sand fine mixture, especially when a large amount of high plastic fine is involved. Silt have to pass #200 sieve. The following strict procedure is followed in mixing so ils containing fines: 1) Determine the weight percentage of dry fines and clean sand. 2) Obtain the water content of each constituent, including silt and sand of required sizes, which were previously sieved and stored in covered containers. 3) Determine the dry w eight of the soil required for preparing ten triaxial samples, usually around 20 pounds. 4) Calculate the dry weight of all constituents required to make 20 pounds of a dry soil mix.

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106 5) Use the water contents obtained in (6) to calculate the required moist wei ght of all constituents. 6) Weight an exact amount of each soil constituent. 7) Mix these soil constituents in a large pan, until it reaches a satisfactory degree of uniformity. The time required to achieve a condition of a uniform mix increases with the amoun t of its plastic fines. Mixing is continued until no visual color variation can be detected. The effort could take as long as a couple of hours, when a large percent of plastic fine is involved. 8) Determine the amount of water to be added to result in a wat er content of 6 9%. A soil mix of a higher fine content tends to require a higher water content to reduce the tamping effort needed to achieve a required compacted density. 9) Weight the required amount of water and spray it intermittently onto the soil whil e mixing. 10) Sieve the soil mix through #8 sieve, whenever necessary, to screen out large moist soil crumbs, which are likely to form, when mixing and wetting a soil containing a large amount of fines. These large crumbs are broken down and remix with the re st of soil to achieve a uniform moisture distribution. 11) Continue (10) until the soil is uniformly mixed with water and no large soil crumbs larger than #8 sieve opening are present. 12) Transfer the moist soil from the mixing pan to a covered container. 13) Store the container with the moist soil in a moisture control room at least overnight to achieve a uniform moisture distribution before it is used to prepare a triaxial test sample.

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107 Specific gravities were determined for the silt and uniform Monterey No. 0/30 sand. The specific gravity of silt is 2.67 and specific gravity of sand is 2.65. The specific gravity of soils containing fines was obtained by mixing calculated amounts of different percentage of fines and uniform Monterey No. 0/30 sand. As an exampl e, the specific gravity of soils containing fines, twenty five percentage of fines and seventy five percentage of uniform Monterey No. 0/30 sand, was calculated as follows: G SF = 5 The weighted specific gravity for soil containing fines was also determined using the same procedure. The dry unit weight at 50% relative density for each parent sand was determined using following equation. Dr=50%= Void ratios corresponding to dry u nit weights at 50% relative density, were determined using the equation: d50% = Sample Preparation Extreme care must be exercised during the complete process of sample preparation. All necessary equipment and supplies are properly cleane d and arranged on the countertop for easy access and are checked for their working conditions. All samples are prepared by following the procedure outlined below. (1) Membrane quality control. Before use, each membrane is carefully examined to see if there is any hole. To detect any hole in a membrane, it is sealed onto the base pedestal at one end and onto the loading cap at the other end with silicon grease and rubber o rings.

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108 The rubber o ring shown in Figure 4.5.The membrane is slowly inflated by filling it with water. After the membrane is inflated with water, the outside surface is dried, so that any water drops seeping through the membrane can be detected. After a membrane is thoroughly checked and found to contain no hole, its inner and outer surface a re then dried to prevent any soil particles from sticking to the surface during sample preparation process. (2) Average membrane thickness. Thickness of the rubber membrane was measured at each end of the membrane using the vernier caliper shown in Figure.4.6. Each membrane must be carefully measured for its thickness. For different measurements are taken, two at the top corners and two at the bottom corners, for its double thickness. An average double thickness is calculated by averaging the four measurements. (3) Obtain two dry porous stones. Moist porous stone may result in sample non uniformity due to capillary if not properly dry; the result may be a non uniform sample. (4) Use the dry porous stone as a template and cut two discs of filter paper to a size slightly smaller than the stone. During the placement of soil, an over size filter paper disc tends to be lifted up by the surrounding tuber membrane and result in a void between the filter paper disc and the stone. On the other hand, if a filter paper disc is too much smaller in diameter than the porous stone, a soil will come in direct contect with the porous stone around the perimeter of the filter paper disc. This may result in the blockade of the porous stone drainage paths. (5) Measure the total height of the cell base, two porous stones, two filter paper discs and loading cap using a precision caliper shown in Figure.4.7. Stack them sequentially and check the final stack for level. Measure the total height from the countertop to nearest

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109 cations. Average these four measurements. This average height is termed an initial height without sample and will be used in subsequent steps to determine the final sample height. (6) Determine the water content (w). the water content of a soil mixed can be de termined using a sample procedure out lined as follows: a. Weight a small amount of a moist soil (w i ) in a pre weighed bowl. b. Beat the soil in a microwave oven for about 10 minutes to remove all of moisture. Determine the final weight (w f ) of the dry soil. The water content is determined as: w = (w i w f ) / w f The water content of a moist soil mix can be maintained relatively constant by sealing the container and then placing it in a 100 percent moisture room. (7) Determine the amount of soil required to achieve a desired void ratio. a. At a desired void ratio, e d , a desired dry unit weight of a soil mix, r d , can be determined as: r d = Where Gs is the weighted specific gravity of the soil mix. b. 2 H/4 Where D is the estimated sample diameter and H is the estimated sample height. D can be obtained by subtracting the double membrane thickness from the average inner diameter of the sample perpendicular directions. The average inner diameter of the mold is obtained by taking the average of two diameter measurements at two perpendicular directions. A caliper is used in this diameter measurement. A proper required sample height H, should allow an easy access to the top porous stone and filter paper. Usually it is pref erred to expose the porous stone slightly above the sample mold. Thus, H depends on the height of the mold, and it should be a little less than the measured height from the bottom of the lower porous stone to the top of the mold. This

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110 height will allow for the easy removal of the stone and the filter paper disc during the final stage of sample preparation. c. The weight of dry soil required is calculated as: W d d × V d. Since moisy tamping method is adopted for compacting moist soil, the required weight of mo ist soil, W wet , is calculated as: W wet = W d × (1+ W) where W is the water content of the moist soil mix. (8) Apply a thin layer of grease a high vacuum silicon as sealant around the side of a base pedestal shown in Figure 4.8. (9) Place a porous stone and then a piece of filter paper disc on top of the base pedestal. (10) Stretch the membrane over the base pedestal so that it adheres uniformly on the side wall of the base pedestal. Carefully smooth out the membrane over the base pedestal and eliminate any air voids along the interface. Make sure that the membrane seats straight up on the base shown in Figure 4.9. (11) Carefully place two o rings in the 0 ging grooves on the base pedestal groves. This will help secure the membrane to the base pedestal and prevent leakag e. (12) Apply a thin layer of grease sealant to the edges of the split mold used to from a sample. (13) Place the mold around the base pedestal (14) Tighten the two halves of the split mold together using pipe clamps shown in Figure 4.10. (15) Wrap the top portion of the membrane over the top of the mold and fold the excess membrane round the outside of the mold to ensure a smooth surface when the vacuum is applied to the mold. This step involves stretching membrane with four fingers, two from

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111 each hand. The rubber membra ne should be stretched equally outward, and wrap around the top of the mold. (16) Apply a vacuum of 10 psi, so that the membrane adhere on the inner wall of the mold. Make sure that there are no wrinkles on the membrane shown in Figure 4.11. (17) Check the top o f the mold for level and measure the distance from the surface of the filter paper disc inside the mold to the top of the mold, H t . (18) The zero raining device was placed within the mold as shown in Figure 4.12 and the sand previously weighed was poured into the zero raining device. By slowly and smoothly lifting up the zero raining device, the sand fell through the screened opening on the end of the zero raining device. (19) Level the sample top, place a piece of a filter paper disc on top of the sample followed b y a porous stone shown in Figure 4.13. (20) Apply a thin layer of grease sealant around the side of the loading cap. (21) Place the loading cap on top of the top porous stone shown in Figure 4.14. Make sure the tube connecting to the top cap is located in such a manner that it will not hinder the chamber assembly. (22) Check the top of the cap for level in two perpendicular directions. The maximum Tilting of the sample top is limited to 0.002 times the diameter of the specimen. (23) Check the sample height against the init ial height without sample as obtained in step (6). The difference has to be very close to the desired sample height, H, to ensure a good density control.

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112 (24) Carefully pull the membrane over the top cap and use two o rings (already placed over the cap an d on the line) to secure the membrane to the loading cap. Be sure to keep the membrane smooth and free of any air bubbles. (25) The vacuum should still be on the mold (26) While the mold still under vacuum, apply vacuum 10 psi to the sample from the top cap. (27) Release the vacuum to the mold. (28) Recheck the level shown in Figure 4.15 and the sample height to see if the applied vacuum causes excessive volume compaction. (29) Carefully pry open and remove the mold and check the sample for any irregularity which could late r cause any problem. Also check to see if sample is vertical. (30) Use the vernier caliper to measure the sample height at four different perpendicular locations on top of the loading cap shown as in Figure 4.16. Average the four measurements as the final heig ht with sample. (31) Use a pi tap shown in Figure 4.17 to measure the sample diameter with membrane. Measurements are taken at three locations: near the top, near the middle and near the bottom of the sample. The average sample diameter with membrane is calcul ated by using the following formula: ¼ × (top diameter + 2× mid height diameter + bottom diameter) The final sample diameter is determined by subtracting the double membrane thickness from the average diameter of the specimen with membrane. Knowing the fi nal specimen diameter and height, the specimen volume is determined. The dry unit weight of the specimen is determined by dividing the specimen dry weight by the specimen volume. The specimen void ratio is then d = w )/(1+e)

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113 (32) Apply a vacuum of 5 psi to the top a container half filled with de air water. Connect the water supply line from this container to the bottom of the sample. The value at the bottom of the sample is controlled in such a way that the water ca n be drawn from the container into the sample and out from the top of the sample at a very slow rate to prevent any significant loss of fine, when the water is observed to flow out from the top of the sample shown in Figure 4.18, procedures (29) and (30) a re repeated to check and record any volume change due to flushing. Attempt is made to achieve a +/ 2% error in void ratio after flushing. (33) between the chamber and the ba se. (34) Clean the interface area between the chamber and the cell base and apply a thin layer of grease sealant. (35) Recheck the interface area to make sure that it is free of any sand particles. (36) Plug an open tube to the top of the triaxial cell chamber to preve nt any pressure building up, carefully place the chamber over the sample, and lock it onto the triaxial cell base using a rim locking band. (37) Lower the loading ram gently to check if the ram and sample loading cap are properly aligned. Then raise the ram an d lock it in place with a piston lock. Before placing the triaxial cell chamber onto the base, it is absolutely critical to check if the loading ram is raised to the highest possible position and securely locked. This is to prevent the loading ram to come in contact with the top cap and severely disturbed the sample. (38) Fill the chamber with water, until water flows out of the top of the chamber through the open tube.

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114 (39) After the chamber is properly filled, close the valve leading to the filling chamber and re move the open tube from the top of the chamber. (40) Close the vacuum line connecting to the top of the sample and allow the water, under 5 psi vacuum, to continue to flow into the bottom of the sample. (41) Apply a effective stress of 5 psi disconnect th e water line under 5 psi vacuum at the bottom of the sample, and connect a water supply line to the bottom of the sample to allow desired water to flow into the sample under atmospheric pressure, until the flow of water ceases. At this moment the vacuum in the sample is completely released. (42) Increase the effective stress to 15 psi and allow the sample to consolidate under 15 psi. (43) Connect the bottom of the sample to beck pressure line. Make sure both the effective stress and back pressure burettes are fille d with water. (44) Connect the four way transducer value to the bottom of the sample. (45) Before mounting the transducer, saturate all lines connecting to the four way transducer valve with water. (46) Saturate the four way transducer valve with water by turning it u pside down and gently tapping the side of the value to help getting rid of trapped air. This four way transducer value is used to measure the magnitude of the effective stress , the back pressure to be applied to the bottom of the sample, and the pore press ure at the top pf the sample using a digital multimeter during cyclic testing. (47) Now simultaneously raise the consolidation and back pressures in such a manner as to maintain a 10psi effective stress in the sample. In order to minimize sample disturbance du e to sudden pressure increase, raise the pressure in a 10 psi increment roughly every 2 3 minutes.

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115 (48) Continue to raise the back pressure until it finally reaches 100 psi. (49) Allow the pressure in the sample to reach an equilibrium and then check the B parameter. B value usually can reach better than 0.99 upon the application of the final increment of back pressure. This proves that the sample saturation procedure is very time effective in achieving a satisfactory degree of saturation. Figure 4.5 Rubber o ring (Jungang Liu, 2019 ) Figure 4.6 Vernier caliper used to measure the thickness of the membrane and diameter of the sand specimen (Jungang Liu, 2019 )

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116 Figure 4.7 Vernier caliper used to measure the height of the test specimen. (Jungang Liu, 201 9 )

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117 Figure 4.8 Looking down on the base of the triaxial cell. (Jungang Liu, 201 9 ) Figure 4.9 Rubber membrane placed on the base of the t riaxial cell. (Jungang Liu, 2019 )

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118 Figure 4.10 cylindrical mold placed on the base of the triaxial cell. (Jungang Liu, 2019 ) Figure4.11 T he membrane adhereon the inner wall of the mold. (Jungang Liu, 2019 )

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119 Figure 4.12 Soil sample placed within the zero raining device. (Jungang Liu, 2 019 ) Figure 4.13 Porous stone placed on top of the Monterey No. 0 Sand Specimen. (Jungang Liu, 2019 )

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120 Figure 4.14 Top cap placed on top of the specimen (Jungang Liu, 2019 ) Figure 4.15 Positioning of the top surface of the loading cap parallel to the base of the triaxial cell (Jungang Liu, 2019 )

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121 Figure 4.16 Measuring the length of the sample using a vernier caliper (Jungang Liu, 2019 )

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122 Figure 4.17 Pi tape (Jungang Liu, 2019 ) Figure 4.18 Specimen was saturating (Jungang Liu, 2019 )

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123 Samples Quality Control Relative Density Control To ensure that representative and accurate results were obtained from the triaxial test, a specified procedure was determined and followed. Relative den sity is expressed as: D r d min max min ) ×100% Where: D r = relative density (in percent) max = maximum dry density (unit weight) of the soil min = minimum dry density (unit weight) of the soil d = in place dry density (unit weight) of the soil To determine the air dry weight of soil necessary to create the sample, the following equation was used: d = w s s 2 ) (H s )/ 4 Whe d = unit dry weight of the sample at desired relative density w s = dry weight of specimen D s = diameter of specimen H s = height of specimen The diameter and height of the specimen are theoretical final measurements and were determined by calculating the mold size and subtracting the membrane thickness. Several trial samples were prepared initially to ensure proper density was being achieved. Soil Sample Saturation Water levels in the burettes connected to the cell and back pressure line were filled to appropriate levels.

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124 To limit the amount of stress the specimen undergoes during the saturation phase, the back pressure and cell pressure were increased simultaneously, maintaining the cell pressure 15psi greater than the back pressure. This procedure was carried out slowly so that the pore pressure throughout the specimen was maintained the equilibrium. In most cases, the back and cell pressur es were simultaneously raised to 75 psi and 90 psi respectively, following which time the specimen was allowed to saturate overnight. Determining B Parameter After saturating for approximately 24 hours, the B Parameter was checked. With back pressure drai nage lines closed, the cell pressure was increased by 0.5mv. By measuring the using the following equation: 3 ter 3 = change in cell pressure, change in consolidation It was desirable to obtain a B parameter of greater than or equal to 0.95 before the specimen was considered satura ted. For triaxial samples in this paper, after saturating for approximately1 2 days, the B parameters were equal to or greater than 0.95. Test Procedure Although degree of saturation is usually satisfactory for testing upon the application of final back pr essure increment, a sample is always allowed to sit overnight before it is tested. A 20 kip MTS electro hydraulic machine is used in performing all cyclic triaxial tests. Procedure for testing a sample is outlined as follows:

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125 (1) Check the pore pressure B par ameter is higher than 0.95. All soil specimens were allowed to consolidate over night at the end of sample preparation, when the B value is found satisfactory. (2) Transfer the triaxial cell to the MTS machine, center and lock the triaxial cell base to the MTS loading platform using three C clamps. (3) Calculate the load required to balance the consolidation force acting on the triaxial cell loading ram. This load is equal to the magnitude of the cell pressure times the cross sectional area of the loading ram. (4) Use MTS machine to apply a slightly larger load than required to balance the consolidation pressure to bring down the loading ram and then connect the loading ram to the sample loading cap. Record any volume change open the completion of the connection . (5) Calculate the deviator load required to produce a deviator stress corresponding to a desired stress ratio. This load is equal to two times the product of initial effective stress and stress ratio times the cross sectional area of the sample. (6) Set the required deviator load level on MTS machine. An invert sine wave form with a frequency of 0.5 Hz was used in this test program. (7) Set the excess pore pressure time and deformation time plotter for proper scales. (8) Zero and set the proper sampling frequency a nd scales for data logger which records digitally the deviator load, the excess pore pressure, and the sample deformation. (9) Close the value connecting to back pressure line and switch the transducer to measure pore pressure from the top of the sample. (10) Open computer on the right side of MTS machine. (11)

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126 (12) Test setup on dialog windows, select cyclic test, control mode is force and channel is (13) disabled, Preset 50 cycles on tees counters. (14) Save data file to my triaxial test file. (15) Zero force and displacem ent on signal auto offset before run test. (16) Test setup on computer shown in Figure 4.19. (17) Start the test and observe the load deformation, excess pore pressure time, and deformation time plotting. (18) Terminate the test when a sample has liquefaction during c yclic loading, or when the excess pore pressure and the axial deformation are stabilized. (19) Allow the excess pore pressure in the sample to dissipate by opening the back pressure valve, and then disconnect the sample loading cap from the loading ram. (20) Remove the triaxial cell from the MTS machine. (21) Remove the tested soil form the triaxial cell for recycling and clean the cell for subsequent use.

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127 Figure 4.19 Set up on the computer to run MTS. (Jungang Liu, 2019 )

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128 Tests Results A series of i sotropically consolidated undrain cyclic triaxial tests were conducted to investigate the effect of fine contents on the liquefaction resistance of soils. All triaxial specimens were prepared to attain 2 inches in diameter and 4 inches in length under thre e different relative densities (30%, 45% and 60%). Two effective stress s of 15psi and 30 psi, three cyclic stress ratios (0.2, 0.3 and 0.4) were used for series of teste on the uniform medium Monterey sand containing fines at half hertz frequency. A total of 96 cyclic triaxial test was performed on the uniform medium Monterey No. 0/30 sand with six different percentages of fine content (5%, 10%, 15%, 25%, 35% and 45%) and plasticity index 20 showed in Table 1. In this chapter, one set of cyclic tr iaxial test result is shown, and it includes soil liquefaction potential curves, cyclic axial load versus number of cycles to liquefaction and excess pore water pressure versus number of cycles to liquefaction. The rest of test results put in the Appendix D. Table 4.1 All soil samples in cyclic triaxial test (Jungang Liu, 2019 ) 45 * : Fine content 45% was only prepared at a relative density of 30 percent. Cyclic Triaxial Test Relative Density (%) PI Fine Content (%) Cyclic Stress Ratio Effective S tress (psi) Frequency (Hz) Total 30 20 5 0.2 15 0.5 96 45 10 0.3 30 60 15 0.4 25 35 45 *

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129 Soil Liquefaction Potential Curves In the Figure 4.20, 4.21, 4.22, 4.23, 4.24 and 4.25, it showed that 96 cyclic triaxial tests conducted by Jungang Liu. In 96 cyclic triaxial tests, another uniform medium sand (Monterey sand) with PI 20% and fine contents were 5%, 10%, 15%, 25%, 35%,45%, w ere prepared under 30% ,45% and 60% of relative densities, and run in 15 psi and 30 psi effective stress s. In the Figure 4.20 and 4.21, it presented that cyclic stress ratio versus number of cycles to liquefaction for soil samples with PI 20% and six diff erent fines contents 5%, 10%, 15%, 25%, 35%,45%, prepared at relative density of 30% and under 15psi,30psi effective stress . In the Figure 4.20 and 4.21, it indicated that number of cycles to liquefaction did not increase with increasing percent of fine c ontent under the cyclic stress ratio , but number of cycles to liquefaction in crease with increasing effective stress . In Figure 4.20 and 4.21, it showed that soil with fine content 15%, prepared at 30% relative den sity and run under 15psi, 30psi effective stress s, had the smallest number of cycles to liquefaction under the same cyclic stress ratio . Soil sample with fine content 45% had the largest number of cycles to liquefaction under the same cyclic stress ratio . In the Figure 4.22 and 4.23, it showed th at cyclic stress ratio versus number of cycles to liquefaction for soil samples with the same PI and five different fines contents 5%, 10%, 15%, 25%, 35%, prepared at relative density of 45% and under 15psi,30psi effective stress . In the Figure 4.22 and 4.23, it also indicated that number of cycles to liquefaction did not increase with increasing percent of fine content under the cyclic stress ratio , however number of cycles to liquefaction increase with increasing effective stress . Compare to Figure 4.20 and 4.21, the bigger relative density caused the more number of cycles to liquefaction. In Figure 4.22 and 4.23, it also showed that soil had the smallest number of cycles to liquefaction under the same

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130 cyclic stress ratio when sample contained fine conte nt 15. Soil sample with fine content 35% had the largest number of cycles to liquefaction under the same cyclic stress ratio . In the Figure 4.24 and 4.25, it presented that cyclic stress ratio versus number of cycles to liquefaction for soil samples with PI 20% and different fines contents 5%, 10%, 15%, 25%, 35%, prepared at relative density of 60% and under 15psi,30psi effective stress . In the Figure 4.24 and 4.25, it also indicated that the similar tests results with different relative densities 30% and 45%. It is number of cycles to liquefaction did not increase with increasing percent of fine content under the cyclic stress ratio and also number of cycles to liquefaction increase with increasing effective stress and relative density. In Figure 4.24 and 4.25, it also showed that soil with fine content 15%, prepared at 60% relative density and run under 15psi, 30psi effective stress s, had the smallest number of cycles to liquefaction under the same cyclic stress ratio . Soil sample with fine content 35% ha d the largest number of cycles to liquefaction under the same cyclic stress ratio . Figure 4.20 Cyclic s tress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with different percent of fines contents at relative density of 30% and effective stress 15psi. (Jungang Liu, 2019 ) . 0 0.1 0.2 0.3 0.4 0.5 0.1 1 10 100 1000 Cyclic Stress Ratio No. Cycles to Liquefaction Dr=30% PI=20 E.S.=15psi ES15FC5 ES15FC10 ES15FC15 ES15FC25 ES15FC35 ES15FC45

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131 Figure 4.21 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with different percent of fines contents at relative density of 30% and effective stress 30psi. (Jungang Liu, 2 019 ) Figure 4.22 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with different percent of fines contents at relative density of 45% and effective stress 15psi. (Jungang Liu, 2019 ) 0 0.1 0.2 0.3 0.4 0.5 0.1 1 10 100 1000 Cyclic Stress Ratio No. Cycles to liquefaction Dr=30% PI=20 C.P.=30psi ES30FC5 ES30FC10 ES30FC15 ES30FC25 ES30FC35 ES30FC45 0 0.1 0.2 0.3 0.4 0.5 1 10 100 1000 Cyclic Stress Ratio No.Cycles to Liquefaction DR45ES15FC5 DR45ES15FC1 0 DE45ES15FC15 DR45ES15FC2 5 DR45ES15FC3 5 Dr=45% PI=20 E.S.=15psi

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132 Figure 4.23 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with different percent of fines contents at relative density of 45% and effective stress 30psi. (Jungang Liu, 2019 ) Figure 4.24 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with different percent of fines contents at relative density of 60% and effective stress 15psi. (Jungang Liu, 2019 ) 0 0.1 0.2 0.3 0.4 0.5 1 10 100 1000 Cyclic Stress Ratio No.Cycles to Liquefaction DR45ES30FC5 DR45ES30FC10 DR45ES30FC15 DR45ES30FC25 DR45ES30FC35 Dr=45% PI=20 E.S.=30psi 0 0.1 0.2 0.3 0.4 0.5 1 10 100 1000 Cyclic Stress Ratio No. Cycles to Liquefaction Dr60FC5ES15 DR60FC10ES15 DR60FC15ES15 DR60FC25ES15 DR60FC35ES15 Dr=60% PI=20 E.S.=15psi

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133 Figure 4.25 Cyclic stress ratio versus number of cycles to liquefaction for Monterey No.0/30 sand with differen t percent of fines contents at relative density of 60% and effective stress 30psi. (Jungang Liu, 2019 ) Cyclic Axial Load versus Number of Cycles to Liquefaction In Cyclic Triaxial test, a cyclic axial load of constant amplitude (24 Ib) was applied on top of soil specimen with a frequency of o.5 Hz. Figure 4.26 shows cyclic load versus number of cycles to liquefaction in Cyclic Triaxial Test. In the first 8 cycles of Fig.4.26, the amplitude of cyclic loading was held constant without noticeable sample defor mation with increasing number of cycles. It means soil still was strong. The amplitude of cyclic loading begun to decreased with increasing number of cycles after 4th cycles. After 9 th cycles of Figure 4.26, the amplitude of cyclic loading rapidly dropped with increasing number of cycles. That means the both sample were liquefied and too soft in the last few cycles. 0 0.1 0.2 0.3 0.4 0.5 1 10 100 1000 Cyclic Stress Ratio No. Cycles to Liquefaction Dr60FC5ES30 DR60FC10ES30 DR60FC15ES30 DR60FC25ES30 DR60FC35ES30 Dr=60% PI=20 C.P.=30psi

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134 Figure 4.26 Cyclic Axial Load versus Number of cycles to liquefact ion in Cyclic Triaxial Test. (Jungang Liu, 2019 ) Excess Pore Water Pressure versus Number of Cycles to Liquefaction Figure 4.27 shows excess pore pressure versus number of cycles in Cyclic Triaxial Test. During the first 8 cycles of cyclic axial load ap plication, the sample showed no noticeable deformation although the pore water pressure built up gradually. However, during the 9th stress cycle, the pore pressure suddenly increased to a value equal to the externally applied effective stress . In fact, the soil had liquefied and the effective stress had been reduced to zero. Over a wide range of strains, the soil could be observed to be in a fluid condition. Excess pore water pressure continues to build up steadily as the number of stress cycles increase, until there is a sudden increase denoting the onset of initial liquefaction. The different values of pore water pressure developed during increases and decreases in deviator stress reflect the influence of the applied stress conditions. -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 0 2 4 6 8 10 12 14 16 18 20 Cyclic Axial Load(Ib) Number of Cycles to Liquefaction Liquefaction happens

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135 Figure 4.27 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. (Jungang Liu, 2019 ) Cyclic Deviator Stress versus Axial Strain The cyclic deviator stress axial strain graph was shown on Figure 4.28. The cyclic deviator stress of constant amplitude (7.5 psi) was applied on the top of soil specimen. In the CTT, the range of axial strain was 0.15 to +0.15 in the first three cycles. It means that the soil sample was no obvious deformation. In the last few cycles of figure 4.28, th e samples turn larger axial strain which was 0.3 to +0.3. The cyclic deviator stress dropped 80 percent. It means that soil sample had liquefied. 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore water pressure (psi) Number of cycles to liquefaction Liquefaction happens

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136 Figure 4.28 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. (Jungang Liu, 2019 ) Str ess Path Q v q. At the beginning of cyclic triaxial test, deviator stress (7.5psi) applied on the sample, and also during first 8 cycles of stress application, the amplitude of cyclic deviator stress kept constant with decreasing the effective mean stress. The amplitude of cyclic deviator stress dropped 10 percent after 8 th cycle in the Figure 4.2 9. Excess pore water pressure increased closed to the value of applied effective stress . It means that the sample turned softer and softer. -10 -8 -6 -4 -2 0 2 4 6 8 10 -0.5 -0.4 -0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4 0.5 (psi) a Liquefaction

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137 Figure 4.29 p versus q stress path in Cyclic Traixial Test. (Jungang Liu, 2019 ) Effect of Fines Content on Liquefaction Resistance It shows in the chapter 7. -10 -8 -6 -4 -2 0 2 4 6 8 10 0 2 4 6 8 10 12 14 16 18 20 q (psi) p' (psi) liquefaction

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138 CHAPTER V HOLLOW CYLINDER TEST APPARATUS, PROGRAMS AND TEST RESULTS ANALYSIS Background A hollow cylindrical apparatus (HCA) is an extremely valuable tool for studying constitutive behavior under generalized stress conditions. The HCA allows independent control of the magnitudes of the three principal stresses and rotation of the major minor principal stress axes under simple shear conditions prevailing in earthquake shaking while recording t he specimen deformational and pore pressure responses. The University of Colorado at Denver Hollow Cylinder Torsional /Axial test cell was designed and fabricated by Dr. Jing Wen Chen while conducting his doctoral research at UC Denver in 1988. In the holl ow cylinder test at UC Denver, a hollow cylindrical soil specimen is enclosed in between an inner membrane and an outer membrane. The effective stress can be independently applied on both inner and outer chambers; therefore, inner and outer pressures can b e controlled either equally or unequally. The axial load and torque are applied on the top of specimen and transmitted by a top cap or a pedestal to the specimen. When each of these boundary stresses can be controlled independently, both the principal stress direction and the relative magnitude of the intermediate principal stress can be controlled, thus the hollow cylindrical test (HCT) can facilitate more generali zed stress path testing than the conventional test apparatus. It is also possible to control (or measure) the pore water pressure and apply back pressure, so that drainage conditions can be controlled and both drained and undrained tests can be performed. As a result, the HCT offers an opportunity of extending the stress path approach to include simulation of both principal stress rotation and variation in

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139 intermediate principal stress, as well as conducting fundamental research into the effect of principal stress rotation under a reasonably generalized stress state . Hollow Cylinder Test Apparatus Principles of Hollow Cylinder Testing Figure 5.1 illustrates idealized stress conditions in a hollow cylindrical element subjected to axial load, W , torque, M T , internal pressure, P i , and external pressure, P o . During shearing, the torque, M T , develops shear stresses, and ( = ) in vertical and horizontal planes the axial load, W , contributes to a vertical stress, z . P i and P o , establish a gradient r , across the cylinder wall. The relationship between radial stress, r , and the circumferential stress, , is expressed by the equilibrium equation: r r / dr ) (5 1) where r is the radial distance to a point in the hollow cylinder, and r and are the radial and circumferential stress increments respectively. The stress condition in an element of a hollow cylinder specimen is shown in Fig. 5.1. Both inner and outer pressure are applied on the membrane so that there is no shear stress on the vertical boundaries, r is always a principal stress because there are no shear stresses on circu mferential surface throughout the wall.

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140 Figure 5.1 Idealized stress and strain components within the HCA subjected to axial load, W , torque, M T , internal pressure, Pi , and external pressure, Po : (a) hollow cylinder coordinates; (b) element component stresses; (c) element component strains; (d) element principal stresses (after Zdravkovic and Jardine, 2001).

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141 (5 2) Since the stresses will not be uniform across the wall of the cylinder for various loading conditions, to consider the hollow cylinder as an element, it becomes necessary to calculate z r . Hight et al. (1983) used the following expressions: z 2 a 2 )] + [(P o b 2 P i a 2 ) / (b 2 a 2 )] (5 3) r = (P o b + P i a) / (b + a) (5 4) Average cir = (P o b P i a) / (b a) (5 5) = 3M T 3 a 3 ) (5 6) r , is usually equal to the intermediate principal stress ( 2 ). The major and minor principal stresses, 1 and 3 , are observed from the average r , and as following: 1 = [( z )/2] + [( z )/2] 2 + ( ) 2 (5 7) 2 = r (5 8) 3 = [( z )/2] [( z )/2] 2 + ( ) 2 (5 9) By regarding the specimen as a single element, the state of strain is presented in cylindrical coordinates in terms of the following components:

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142 (5 10) Also, it is necessary to calculate the average strains. According to the paper of Hight et al. (1983), the average strains are calculated using the following equations: Average axial strain z (5 11) Average radial strain r = [(u o u i )/ (b a)] (5 12) Average circumferential strain = [(u o +u i )/ (b+a )] (5 13) Average shear strain o 3 r i 3 )]/ [3H (r o 2 r i 2 )] (5 14) Where the definitions of average stresses and strains are shown in Figure 5.2. Since the average va lues of z and are based on strain compatibility only, the expressions for the average strains are valid and independent of the constitutive law of the material. The average values of r and are based on a linear variation of radial displacement across the wall of the specimen. In the hollow cylinder test, the radial strain ( r ) is usually the intermediate principal strain, 2 . The major and minor principal strains can be observed from the average strain components: 1 = [( z )/2] + [( z )/2] 2 + [ /2 ] 2 (5 15) 2 = r (5 16) 3 = [( z )/2] [( z )/2] 2 + [ /2] 2 (5 17)

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143 Parameters and b are two variables of stress path to describe fundamentally different aspects in the applied state of stress. (as shown in Figure 5.1(d)), is the in clination of major principal stress direction with respect to the vertical axis, which can be varied from 0 to 90°. It can be computed from the known average stress components tan 2 = 2 r ) (5 18) b is defined as the relative magnitude of the intermediate principal stress, which can be varied from 0 to 1: 2 3 1 3 ) (5 19) For the particular case of equal internal and external pressure, P i =P o =P , and are usually assumed to be equal to P r = (P o r o + P i r i ) / (r o + r i ), 2 is equal to P as well. Therefore, changes in the angle are accompanied by changes in magnitude of b . When Pi=P o b = sin 2 et al., 1983) (5 20) The direction of strain increment can be calculated from the incremental strain components tan2 = d / (d z d ) (5 21) The amount of non coaxiality was defi ned as the difference between the directions of principal stress and of principal strain increments as, . Stress Distribution in Hollow Cylinder Specimens The most critical aspect of the use of hollow cylinder specimen is the nonhomogeneity of stress and strain distributions, developed in the wall of a specimen as a result of curvature of the wall and end restraint. Stress nonuniformity due to curvature can be minimized by selecting an appropriate geometry of the specimen.

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144 Even though hollo w cylinder devices offer highly promising capabilities for the study of soil behavior , their use has been subjected of criticism. These objections arise principally due to the non uniform distribution of stresses and strains within the specimens. Stress no n uniformities occur across the wall of a hollow cylinder due to the specimen geometry, end restraint, the application of torque or different internal and external pressures. The tested specimen size affects significantly the stress non uniformity level. W hen the wall thickness is reduced or the inner radius is increased, the stress distribution becomes more uniform (Sayao and Vaid, 1991). Because it is not easy to measure either the stresses or the strains across the wall of the hollow cylinder directly, it becomes essential to set bounds to the differences between the calculated and real averages and the magnitude of deviations from the real averages. By using the finite element method and assuming that material behaves as either isotropic or elasto plast ic (modified Cam clay), Hight et al . (1983) defined the non uniformity coefficients 1 and 3 for individual stress components, as shown in Figure 5.3. The magnitude of the difference between calculated and real stress average can be characterized by norma lized parameter 1 : 1 = (5 22) where is the real average, L r )/2], is a measure of the stress level. Therefore 1 is inversely related to accuracy. 3 is the parameter to quantify the level of non uniformity of stresses: 3 = (5 23) where is the distribut ion of the particular stress, z or under consideration across the hollow cylinder specimen. 3 may be used to minimize the difference between the actual stress distribution and the real average.

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145 For differences in strain averages and strain non uniformities, similar definitions for 1 and 3 are used. According to Hight et al. (1983), the magnitudes of 1 and 3 are dependent on stress state, specimen geometry and the constitutive law of the speci recommended keeping stresses within a limit where the ratio of outer to inner cell pressures is 0.9


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146 between the stress strain and strength response of hollo w cylinder simulations and a uniform single element. Rolo (2003) used a classical elasto plastic non linear, modified Cam clay soil model with a finite element method to analyze most of the features that were thought to influence the development and magni tude of non uniformities. The non uniformity increased as the specimen approached the failure surface, which agreed with the observations by Hight et al. (1983) on specimens with fixed ends. The specimen with free ends resulted in more uniform conditions. The results revealed that non uniformities could result in either over or underestimation of certain stress and strain parameters. Figure 5.2 Definitions of average stresses and strains (after Hight et al., 1983)

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147 Figure 5.3 Definitions used for stress non uniformity and accuracy (after Hight et al., 1983). Specimen G eometry The uniformity of the stress distribution across the wall of hollow cylinder specimens is affected by the specimen geometry, both the curvature and end restraint. This result came from the detailed study of stress distributions using both isotropic linear elastic and plastic formulations to represent the soil in specimens of different geometries under different load combinations. A suitable height of the specimen can engender reasonably uniform distributions of stress (Hight et al., 1983). The differences b etween real and calculated averages of stress and strain were attributed to the selected specimen geometry and the stress path. As the ratio of inner to outer radii, r i /r o , approaches unity, both 1 and 3 reduce. Figure 5.5 was produced by Porovic (1995) by assuming a linear variation of applied shear stresses, , and a linear elastic constitutive law, to display the ratio of maximum and minimum shear stresses to average shear

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148 stress for three different specimen dimensions. As the diagram shows, the leve l of non uniformity for a fixed wall thickness would reduce with the increase of specimen diameter. Therefore, the degree of the stress difference between the calculated and real average was minimized as the inner radius of specimen increased. The selectio n of a suitable geometry for the hollow cylinder specimen would reduce stress non uniformities to an acceptable level. Saada (1988) also quoted that selecting particular specimen geometry played a major role in reducing non uniformity of stress distributio n. Firstly, for sand specimens, an appropriate wall thickness should be applied to meet the following criteria: a) A wall thickness sufficiently large enough relative to the maximum grain size of the tested specimen so the failure mechanisms would not be constrained. b) A specimen volume sufficiently large in relation to the potential volume change resulting from membrane penetration. c) A uniform density across the wall. In order to determine a reasonable specimen geometry, based on elasticity theory a nd the assumption that the central zone, free from end effects should be the same length as the zone influenced by the platens, Saada and Townsend (1981) suggested the following criteria for the specimen geometry: a) (r o r i ) b) Inner radius r i : n= (r i / r o where H is the height, r i and r o are the inner and outer radii of the specimen, and n is the ratio of inner and outer radii. The criteria proposed by Sayao and Vaid (1991) were as follows:

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149 a) Wall thickness: r o r i = 20 to 60mm b) Inner i / r o c) o Membrane Penetration Errors In the hollow cylinder test, rubber membranes are used to enclose the specimens. The effect of membrane penetration on the external measurement of volumetric deformations is attributed to the flexible membrane penetrating into or withdrawing out of the ext ernal voids of volume change in a drained test, and the magnitude of the pore water pressure measured in an undrained test. Therefore this effect should be accounted for to make a confident assessment of actual stress strain behaviour of saturated granular materials in a test. For materials of medium sand size having mean particle size of D 50 specimens, correction for the membra ne penetration is of great importance and should be applied (Molenkamp and Luger, 1981). Studies of the effect of membrane penetration have been undertaken and the particle size of the material is identified to be the major factor to influence the membrane penetration (Frydman et al ., 1973). Theoretical expressions for the unit membrane penetration suggested by Baldi and Nova (1984) and Kramer and Sivaneswaran (1989) are as following: A MP MP = ½ d/D V soil h d)/(E m t m )] 1/3 (5 25) MP = 0.395d(1 2 4 )] 1/3 h d)/(E m t m )] 1/3 (5 26) where MP = unit membrane penetration (in mm); A MP = surface area of membrane (in mm); d = mean particle size, D 50 (in mm); D= Specimen diameter (in mm); V soil = volume of soil

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150 specimen (in mm3); E m t m = thickness of membrane (in mm); h = effective stress (in kPa). A new approach for the assessment of MP was obtained from the differences between measured volume strain of the specimen and the volume of the inner chamber using a single hollow cylindrical specimen under hydrostatic loading by Sivathayalan and Vaid (1998). The proposed expression for the unit membrane pen etration is: m = (5 27) where m is the unit membrane penetration; sr and ir are the measured volume changes of the inner chamber and the specimen, respectively; n is the ratio of the outer to inner radii of the specimen, and A im and A om are the surface areas of the specimen covered by the inner and outer membranes, respectively. Kuwano (1999) evaluated the apparent volumetric strains due to MP over the vertical sides of the specimens using Ham River Sand specimens with rough and lubricated ends. By comparing the measured volume deformations with a conventional volume gauge and wit h local instrumentation, she obtained the following relationship for MP based on isotropic loading/unloading/reloading tests: MP = C MP h = C MP h / h0 ) (5 28) where C MP is a parameter that depends on specimen size and density, membrane thickness and elastic modulus, and on particle shape and size; h and h0 are the current and initial effective stress C MP is 0.015mm for 100mm diameter speci mens of Ham River Sand encased in a 0.5mm thick latex membrane. Kuwano (1999) found that Eq. (5 27) matched the expressions suggested by Baldi and Nova (1984) and Kramer and Sivaneswaran (1989) very well.

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151 Figure 5.4 Effect of stress ratio level on non uniformity coefficients (after Vaid et al. , 1990)

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152 Figure 5.5 Shear stress distribution in hollow cylinder torsional shear test specimens (after Porovic, 1995).

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153 Soil Samples Preparation Soil Samples The same soil samples were using in Chapter 4.4.1. Mixing Soil Samples The mixing soils procedure is the same as in Chapter 4.4.2. Samples Preparation Preparation of a hollow cylinder sand specimen consisted of the following steps. (1) The thickness of the rubber membrane along the axial direction is nonuniform, therefore the membrane thickness is measured only along the areas where the sample is located. The measurements are obtained at the top, middle and bottom of the sample in orthog onal directions. (2) Assemble the inner mold pieces with the balloon inside. Place the aluminum end caps on the top and bottom of the inner mold pieces. Inflate the inner balloon to around 10 psi. At the beginning of the assembly process, the inner mold is in an inverted position. After step (6), the assembly is inverted to the upright position. (3) Wrap the inner membrane around the already assembled inner mold. (4) Place the plastic stand around the inner mold assembly. (5) Slip the bottom pedestal into the inner mold a nd partially deflate the balloon that is inside the inner mold. Once the balloon is partially deflated, the bottom pedestal will slip down to rest on top of the plastic stand. (6) Apply a thin layer of vacuum grease around the outer edge of the bottom sample p edestal. Wrap the membrane around the edge and seal the membrane using several O rings.

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154 (7) Invert the inner mold and plastic stand assembly, and place the assembly on top of the bottom plate. Bolt the bottom pedestal to the bottom plate using 6 Allen head bol ts. (8) Remove the plastic stand from the assembly. Place a piece of ring shaped membrane on the top of the Allen head bolts. This piece of ring shaped membrane is a seal to prevent vacuum leakage when a vacuum is applied between the outer membrane and outer mold. (9) Apply a thin layer of vacuum grease around the outer edge of the exposed top of the bottom sample pedestal. Place the outer mem brane over the exposed outer ed ge of the bottom sample pedestal and seal the membrane using several O rings . (10) Place the outer mold over the inner assembly and on top of the ring shaped membrane. Place a thick layer of vacuum grease around the bottom outer edge of the outer mold to seal the gap between the outer mold and the ring shaped membrane. (11) Wrap the outer membrane arou nd the upper end of the outer mold. Seal the membrane to the upper end of the outer mold using one O ring. (12) Apply a vacuum to the vacuum ports located on the wall of the outer mold. This will create a vacuum in the space between the inner wall of the outer mold and the outer membrane. Shape the outer membrane to be free of any wrinkles. (13) Weigh the desired amount of soil to be used in the sample. (14) Using the long neck funnel, uniformly deposit the sample sand in the space between the outer membrane and the inne r mold in a consistent fashion. A uniform deposit is achieved by placing the mouth of the long neck funnel at the top of the sand deposit so that there is no distance for the sand to drop. (15) Smooth the surface of the sand with a wooden plate. It is very impo rtant that the top of the sand sample is smooth to prevent necking of the sample.

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155 (16) Place the top sample pedestal on top of the wet surface of the sample sand. Press the top sample pedestal into the sand until the stainless steel torsion plates are fully emb edded into the sample. Use a bubble level and a height caliper to obtain a sample of uniform height. (17) Apply a thin layer of vacuum grease around the outer edge of the top sample pedestal. Wrap the outer membrane around the top sample pedestal and seal the m embrane using several O rings. (18) Apply a thin layer of vacuum grease around the annular of the top sample pedestal. Wrap the inner membrane around the top sample pedestal as seal the membrane using several O rings. (19) Disconnect the vacuum from the ports on the outer wall of the outer mold. Slowly apply vacuum to the top of the sample. A 10 psi vacuum was applied to the sample used in this study. (20) Remove the outer mold from the sample assembly. Deflate the balloon from the inner mold, and remove the inner mold, p iece by piece. The aluminum bottom cap of the inner mold cannot be removed from the top of the sample. (21) To remove the aluminum bottom cap, loosen the bottom platen from the base plate. Remove the inner mold aluminum bottom cap and replace the bottom platen. Tighten the six allen head bolts connecting the bottom platen to the base plate. (22) Supply deaired water to the bottom of the sample. Flow rate and seepage pressure must be controlled to prevent sample disturbance. A total of 2 psi pressure was applied in th e preparation of all of the samples used in this study.

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156 (23) Measure the outer diameter of the sample at the top, middle and bottom of the sample. A top, middle and bottom of the sample using the membrane thickness values obtained earlier. Record the top, middle and bottom sample diameters. (24) Measure the inner dimension of the sample using an inside micrometer. Measure points at top, middle and bottom of the sample at a spread of 0 o , 45 o , 90 o and 135 o from the initial orientation. This will give 12 points of measurement. Add the average membrane thickness readings obtained at the top, middle and bottom of the inner membrane, to the readings obtained above. This will give the top, middle and bottom inner sample diameter. (25) Measure the height of the sample from the bottom of the bottom plate, to the top of top sample pedestal. Obtain readings at three equally spaced points, using a height caliper. Subtract the height of the t op sample pedestal portion that the top cap will rest on, from the overall height dimension. Subtract the bottom sample pedestal portion that rests on the bottom plate. This will give the average final height of the sand sample. (26) Install the four stainless steel support bars on the base plate and bolt them to the plate using four a llen bead bolts. (27) Connect the piston and top cap with the tapered pin. Fasten the pin using the spacer and nut. (28) Slip the top plate over the piston assembly a short distance, and tig hten the locking mechanism to lock the piston. (29) Place the top plate and piston assembly on the top of the four stainless steel support bars. Bolt the plate to the bars using the four allen head bolts.

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157 (30) Loosen the locking mechanism on the piston and carefully lower the piston and top cap assembly to rest on the top sample pedestal. Do not disturb the sample during this procedure. (31) Tighten and piston locking mechanism and connect the top cap and top sample pedestal using the sic Allen head bolts. When torque is applied to the Allen head bolts, the locking mechanism will absorb the torque reaction, not the sample. (32) Using a forklift machine, lower the entire assembly below the hanging chamber cell. Center the assembly beneath the cell and raise the assembly into the chamber cell. (33) Using three large bar clamps, clamps the chamber cell to the assembly. (34) Shift the chamber cell stoppers located on the top plate, into the proper positions. Tighten the stoppers in the lock position to prevent uplift of the chamber cell dur ing the test. (35) Carefully mount the entire assembly in the MTS machine. Mount the assembly on top of the load cell plate and connect the assembly to the load cell plate using four allen head blots. (36) Supply water to the inner and outer chamber of the sample si multaneously. (37) Gradually apply a effective stress increment of 2 to 3 psi while lowering the vacuum by the same increment, maintaining the initial effective stress placed on the sample by the vacuum (38) Continue to decrease the vacuum in increments until zero vacuum. The effective stress is now due entirely effective stress . (39) Adjust the effective stress and back pressure simultaneously until the desired values are obtained.

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158 Samples Quality Control Relative Density Control To ensure that representative and accurate results were obtained from the triaxial test, a specified procedure was determined and followed. Relative density is expressed as: D r = x x100% (5.29) Where: D r = relative density (in percent) max = maximum dry density (unit weight) of the soil min = minimum dry density (unit weight) of the soil d = in place dry density (unit weigh t) of the soil To determine the air dry weight of soil necessary to create the sample, the following equation was used: d = (5.30) d = unit dry weight of the sample at desired relative density w s = dry weight of specimen D outer = outer diameter of specimen D inner = inner diameter of specimen H s = height of specimen The diameter and height of the specimen are theoretical final measurements and were determined by calculating the mold size and subtracting the membrane thickness. Several trial samples were prepared initially to ensure proper density was being achieved.

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159 Soil Samp le Saturation The procedure of soil sample saturation is the same as in Chapter 4.3.2.2. Determining B Parameter The method of determining B Parameter is same as in Chapter 4.3.2.3. Tests Program General Considerations The evolution of laboratory apparatuses for the investigation of stress strain behavior of soils in general was presented by Chen (1988). A focused literature review is presented herein on recent research with hollow cylinder soil testing devices for cycli c soil behavior investigation, with particular emphasis on undrianed pore pressure response and cyclic strength. The fundamental purpose for hollow cylinder soil tests (particularly when torsional loading is available) is the exercise of control, independ ently if desired, over principal and deviatoric stress direction and magnitudes that might be useful in replicating field behavior or calibrating constitutive models. A hollow cylinder of soil is constructed by either: (1) placement, remolded, by some mean s between two flexible, impermeable membranes (inner and outer) temporarily restrained in the proper shape against a rigid form, or mold; or (2) trimmed from a larger, intendedly undisturbed sample of in situ material and wrapped by inner and outer imperme able membranes. The membrane encapsulated specimen, when supported by effective stress due to external pressure or internal vacuum or both resembles a pipe made of soil. Consolidation fluid pressure may be independently applied to the inner and outer memb ranes to vary the distribution of total static stresses within the specimen. Pore fluid (usually water) may pass into or out of the specimen through the ends of the cylinder; pressure

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160 within the specimen voids is varied and usually measured through the end s as well to regulate global effective stress. Porous, flat rings of stone or other rigid material are usually employed as filters at the ends to prevent solids from washing out with fluid flow. Axial loads or torque or both are applied to an end cap or pe destal at one end or the other of the cylindrical specimen and are transmitted to the confined soil through the porous elements. Development of Specifications The UCD hollow cylinder test apparatus was designed in accordance with the conditions listed earl ier as established by Lade (1981) and within the constraints imposed by five aspects common to all laboratory soil testing: specimen dimensions, instrumentation placement and purposes, anticipated specimen strengths, loading machine capacity, and specimen preparation considerations. The UCD apparatus includes features to accommodate construction of soil specimens within the device, membrane encapsulation, and drainage provisions. The ultimate size of the UCD apparatus was dictated by specimen dimensions. T he platens and membrane retention components of the device were built to accommodate preparation and testing of 10inch tall hollow cylinders of soil with an inside diameter of 8 inches and an outside diameter of 10 inches. These dimensions were shown by Hi ght, Gens and Symes (1983) to produce essentially uniform stress and strain distributions across specimen wall thickness and with height, particularly along the central 5 inches of length. This central section should be region within which test parameters are measured for detailed stress strain analysis; overall specimen strengths were investigated in the present study as measured at the specimen ends, and should not be strongly affected by internal non uniformities in an apparatus with the above dimensions .

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161 Test measurement instrumentation versatility was considered in the design of interior space for the pressure chamber and its supporting framework. Future researchers using the UCD apparatus may desire to install deformation and stress monitoring devices within the hollow center of the soil specimen, within the specimen itself, or within the annulus between the specimen and the pressure chamber. Sufficient working space is available to these purposes. The apparatus is ported to allow independent external control of fluid flow or pressure measurement to or from the specimen and the consolidation fluid spaces. The UCD apparatus was conceived to subject soil specimens to both axial and torsional loading, depending on the load frame into which the chamber is i nstalled for test. To this end, external loading specifications were back calculated from axial compression and shear resistances typical of an arbitrary test soil. A cohesionless soil with an effective internal friction angle of 35 o , if subjected to an as sumed maximum external effective stress of 250 psi and a back pressure of 50 psi could be expected to resist up to 740 psi axial compressive stress, as same spe cimen would fail in pure shear at 115psi maximum applied shear stress. Chen (1988) assumed a loading piston diameter of 2 inches to calculate the maximum vertical load required to fail such a specimen as follows: (740.0 200.0) × [(10) 2 (8.0) 2 250.0× (2) 2 31) The loader must apply the failure stress in addition to the uplift on loading piston caused by effective stress , as represented by the second term in Equation 6 30. ) max ) avg ) avg = n (1 n)/(1+n 2 ) (5 32) Torque required to fail the same specimen was back calculated by adapting Equation 6 6 , substituting design dimensions and the maximum shear stress and rewriting:

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162 M t z × (r o 3 r i 3 3 4 3 )]/3 = 14.692 in. Ib (5 33) The test apparatus was designed to accommodate, in consideration of the above calculations, at least 20,000 Ib vertical load and 20,000 in Ib torque. Specifications developed for the purchase of the loader used in the current research were based on these computations. The loader, an Instron TM Model 1322 machine with 50,000 Ib axial (compressive or tensile) and 25,000 in Ib torsional capacity. The apparatus was designed to perm it preparation of test specimens directly on the lower pedestal, in the same fashion as conventional triaxial specimens. The relatively thick cylindrical wall of the design specimens, 1 inch, provided adequate working room for preparation of uniform recons tituted specimens. Parts and their Functions The new hollow axial torsional cylinder apparatus as shown in Figure.5.8 consists of eleven detachable components and three accessories used for sample preparation. The bottom platen as shown in Figure.5.9 is co nstructed of aluminum and is used to seal the bottom of the inner chamber. The platen is twelve inches in diameter to match the twelve inch diameter of the MTS load cell platen to avoid apparatus alignment problems. The top of the cylindrical platen has a counter bore to align and constrain the inner mold from movement during sample preparation. Instrumentation could be placed on top of the platen in the inner chamber area for future improvements. There are six holes in the outer rim of the platen for faste ning the platen to the hollow cylinder base plate. Four threaded holes fasten the bottom platen to the MTS load cell. An o ring on top of the bottom platen seals the platen to the cylinder base plate. The bottom platen has one drainage path to supply water to the inner chamber of the sample.

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163 The hollow cylinder base plate as shown in Fig. 5.10 is constructed of aluminum and has three tiers. The outside diameter is nineteen inches, the intermediate annulus is sixteen inches and the inner diameter is twelve inches. The center bore is nine inches in diameter. The outer annulus has four drainage paths that allow for connection to the inner chamber, outer chamber, top of sample and bottom of sample. These holes are drilled laterally into the base plate and allow for the supply of water to the inner and outer chambers, and measurement of pore pressure or volume change of a sample during testing. The exterior of the second tier has a built in O ring to seal the pressure chamber to the cylinder base plate. There are four holes in the top of the second tier to install support bars to connect the top plate to the base plate. In the future instrumentation could be placed in between the support bars and between the inner and outer chambers. The inner tier has six threade d holes on the top to connect the bottom sample pedestal to the base plate, and six threaded holes on the bottom to connect the bottom platen to the cylinder base plate. An O ring on top of the inner tier seals the sample pedestal from the cylinder base pl ate. The bottom sample pedestal is constructed of aluminum and has two tiers on top and three tiers on the bottom as shown in the left hand side of Fig.5.11. The outside diameter of the pedestal is twelve inches. The top annulus is nine inches in diamet er. The bottom intermediate annulus is eight and one half inches in diameter and the bottom inner annulus is eight and three eighth inches in diameter. The inner bore is eight inches in diameter. The one inch thickness of the top tier exactly matches the t hickness of the sample tested. There are six holes bored in the top of the outer annulus to connect the bottom pedestal to the cylinder base plate. The exterior of the bottom annulus is grooved for O rings and provides the seal for the inner chamber membra ne. The exterior of the top annulus is grooved for O rings and provides the seal for the

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164 outer chamber membrane. The top of the outer annulus has a drainage path to connect to the sample being tested. This path is drilled laterally into the base plate and connected to the sample to measure pore pressure or volume change of the sample during testing. A semicircle grove one sixteenth inch deep is circumscribed along the top of the sample pedestal and is connected to the drainage path. Twelve one inch wide, on e quarter inch thick pieces of bronze porous stone are bolted to the top of the top tier, above the circumscribed groove. The porous stones provide filtration to prevent migration of the soil grains. Twelve stainless steel plates one sixteenth inch thick b y one inch wide separate each of the twelve stones and protrude one quarter inch above the stones. These blades counter torque applied to the top of the sample. The top sample pedestal is constructed of aluminum and is similar to the bottom sample pedesta l in design as shown in the right hand side of Fig.5.11. The top of the sample pedestal has two tiers and the bottom of the sample pedestal has two tiers. The outside diameter is eleven and one half inches. The top annulus has an outside diameter of eight and three eight inches. The bottom annulus is one inch thick and has an outside diameter of ten inches. The thickness exactly matches the sample thickness. There are six threaded holes bored in the outer annulus to connect the top sample pedestal to the to p cap. The exterior of the top annulus is grooved for O rings to provide the seal for the inner chamber membrane. The exterior of the bottom annulus is grooved for O rings to provide the seal for the outer chamber membrane. The outer annulus has a drainage path connecting to the top of sample. This path is drilled laterally into the top platen and is connected to the sample to measure its pore pressure or volume change during testing. A semi circular groove of one sixteenth inch deep is circumscribed along the bottom of the top sample pedestal and is connected to the drainage path. Twelve one inch wide one quarter inch thick pieces of bronze porous stone are bolted to the top of the top tier, above the circumscribed

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165 groove. The porous stones provide filtrati on to prevent migration of the sample media. Twelve stainless steel plates one sixteenth inch thick by one inch wide separate each of the twelve stones, and protrude one quarter inch above the stones. These blades transmit applied torque to the top of the sample. The top cap is constructed of aluminum as shown in Fig.5.12. The top cap is connected to the loading piston and is used to transmit the applied loads to the top sample pedestal. The top cap also serves to separate the inner and outer chambers. The outside diameter is eleven and one half inches, and the interior counter bore is nine inches in diameter. The outer perimeter has six vertical holes bored in it to connect the top cap to the top sample pedestal. The top of the top cap has a three inch tall socket with a four inch outside diameter and a two inch interior counter bore. The middle of the socket has a laterally bored tapered hole drilled through both sides to provide for connection to the loading piston. The top cap has two one half inch diamet er holes that are used to equalize the inner and outer chamber pressure. If measurement of volume change in the inner chamber is required, each of the two holes in the top cap can be connected to one of the two drainage paths located in the top plate. One drainage path can be pressurized and the other path can be used to measure the volume change. The top plate as shown in Fig.5.12 is constructed of aluminum and is fifteen and three quarter inches in diameter. The top plate has four holes bored to fasten th e support bars connected to the cylinder base plate to the top plate. The exterior of the top plate has a built in O ring to seal the outer chamber cell to the top plate. A four inch diameter threaded hole is centered in the top plate and is used to attach the top plate to the piston bushing sleeve. Four locking tabs constructed of aluminum are bolted to the exterior portion of the top plate to prevent slippage of the exterior pressure chamber. Two one half inch drainage holes are bored in the top plate. Th ese

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166 holes function as an air bleed to allow cyclic pressurization during the cyclic test. When cyclic testing is conducted to the inner chamber and the other is connected to the outer chamber. The piston bushing sleeve is constructed of aluminum and is nine inches tall and has a five inch outside diameter. The bottom end of the piston bushing sleeve is threaded to attach the sleeve to the top plate. Once attached, these two pieces form a permane nt integral part of the cell and are not separated. Two stainless steel Thompson A 324864 SP bushings are housed in the center of the sleeve. These bushings have a two inch interior diameter to match the diameter of the loading piston. A double lip wiper r ing is installed in the bottom of the sleeve to seal the pressure leakage and prevent contamination of the Thompson bushing from the loading piston. The top of the sleeve has an annulus plate five inches in diameter used to prevent slippage of the Thompson bushing. A piston lock mechanism as shown in Fig.5.13 is constructed of stainless steel is connected to the top plate of the sleeve to prevent unintentional piston movement. The lock mechanism consists of two halves of a ring with lateral bolts that when tightened will prevent up and down motion of the piston. One curved slot is bored in the center portion of each half and is used to bolt the lock to the top sleeve plate. Torsional rotation of the load piston is prevented when the lock is bolted to the to p sleeve plate. The loading piston is constructed of 440 C stainless steel and is twenty inches tall and slightly less than two inches in diameter with a finely polished surface finish. A one half inch diameter tapered bore is located at the bottom of the piston. A tapered pin that is threaded on the small end is used to connect the piston to the top cap. A nut and washer is used on the small end to produce a movement free connection.

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167 Four support bars as shown in Fig. 5.14 are constructed of 440 C stainl ess steel. These bars are twenty three and three quarter inches tall with a one inch diameter and are used to connect the cylinder bottom plate to the top plate. Each end of the support bar is counter bored and tapped to allow bolts to secure the bars to t he plates. Both ends of the support bars have o rings to prevent leakage from the outer chamber. The pressure chamber as shown in Fig. 5.15 is constructed of aluminum and cast acrylic. The total length is twenty five and one half inches with an outer diam eter of eighteen inches and interior diameter of fifteen and one half inches. Each end of the cast acrylic chamber is threaded to fasten aluminum end rings. Each of the rings is four inches tall. The bottom aluminum end ring has an exterior and interior di ameter equal to the cast acrylic, and has a threaded counter bore to allow the ring to attach to the cast acrylic. An o ring is placed between the ring and acrylic to prevent any leakage. The top ring is similar to the bottom ring, but the top two inches h as a smaller interior diameter. This is done to enable the pressure chamber to slip over the top plate and cylinder base plate without contacting the o ring until the final two inches of movement.

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168 Figure 5.6 ertical stress in HCTA 88 (Chen, 1988) Figure 5.7 88 (Chen, 1988)

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169 Figure 5.8 Schematic diagram of the HCTA 88 at the University of Colorado at Denver Geotechnical Laboratory (Chen, 1988)

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170 Figure 5.9 Bottom Platen (Jungang Liu, 2019 ) Figure 5.10 Base Plate (Jungang Liu, 2019 )

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171 Figure 5.11 Bottom Sample Pedestal (left) and Top Sample Pedestal (Jungang Liu, 2019 ) Figure 5.12 Top Cap and Top Plate (Jungang Liu, 2019 )

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172 Figure 5.13 Piston Locking Mechanism (Jungang Liu, 2019 ) Figure 5.14 Supporting Bars (Jungang Liu, 2019 )

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173 Figure 5.15 Pressure Chamber (Jungang Liu, 2019 )

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174 Test Procedure After the application of final back pressure increment, soil sample is fully saturated. An Instron TM Model 1322 machine with 50,000 Ib axial (compressive or tensile) and 25,000 in Ib torsional capacity is used in performing all cyclic hollow cylinder tests . Procedure for testing a sample is outlined as follows: (1) Check the pore pressure B parameter is higher than 0.95. All soil specimens were allowed to consolidate over night at the end of sample preparation, when the B value is found satisfactory. (2) Transf er the hollow cylinder cell to the Instron TM machine, center and lock the triaxial cell base to the Instron TM loading platform using three bolts. (3) Calculate the cyclic torsional torque required to produce a cyclic shear stress corresponding to a cyclic stress ratio. This load is equal to the product of initial effective stress and stress ratio times the cross sectional area of the sample. (4) Set the required cyclic torsional torque level on Instron TM machine. An invert sine wave form with a frequency of 0.5 Hz was used in this test program. (5) Set the transducer of excess pore pressure to the data logger. (6) Close the value connecting to back pressure line and switch the transducer to measure pore pressure from the top of the sample. (7) Open computer on the right side of Instron TM machine. (8) software in MTS Systems. (9) Procedure fi

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175 128 ft.Ibf. (10) Save data file to hollow cylinder test file. (11) Zero axial force, axial displacement, torsional torque and torsional angle on signal auto offset before run test. (12) Test setup on computer shown in Figure 5.16. (13) Start the test and o bserve the torsional torque deformation, excess pore pressure time, and deformation time plotting. (14) Terminate the test when a sample has liquefaction during cyclic torsional torque loading, or when the excess pore pressure and the axial deformation are sta bilized. (15) Allow the excess pore pressure in the sample to dissipate by opening the back pressure valve, and then disconnect the sample loading cap from the loading ram. (16) Remove the hollow cylinder cell from the Instron TM machine. (17) Remove the tested soil fo rm the cell for recycling and clean the cell for subsequent use. Figure 5.16 : Hollow cylinder test setup in software of MTS. (Jungang Liu, 2019 )

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17 6 Analysis of Results A series of cyclic torsional hollow cylinder test were performed to compare of results of liquefaction resistance of soil from cyclic triaxial and cyclic hollow cylinder tests and also investigate effect of fine contents on the liquefaction resistance of soils. Twenty cyclic torsional hollow cylinder tests performed on the uniform medium Mon terey sand with six different percentages of fine content (5%, 10%, 15%, 25%, 35% and 45%) and plasticity index 20. Table 5.1 detailed test program with information including specimen relative densities, effective stress , cyclic stress ratio and frequency. All hollow cylinder samples are inner diameter of 8 inches, outer diameter of 10 inches, and height of 10 inches. The properties of all soil sample are the same as in cyclic triaxial test in Chapter 4. In this chapter, one set of cyclic hollow cylinder t est result is shown, and it includes cyclic torque versus number of cycles to liquefaction and excess pore water pressure versus number of cycles to liquefaction. The rest of test results put in the Appendix D. Table 5.1 Test program details in hollow cy linder test (Jungang Liu, 2019 ) *: each percentage of fine content run under corresponding cyclic stress ratio. Hollow Cylinder Test Relative Density (%) PI Fine Content (%) Cyclic Stress Ratio E ffective Stress (psi) Frequency (Hz) Total 30 20 5 0.4* 15 0.5 20 60 10 0.3* 30 15 0.2* 25 0.3* 35 0.4*

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177 Cyclic Torque versus Number of Cycles to liquefaction In HCT test, a cyclic torsion load of constant amplitude (40Ibf ft) was applied with a frequency of 0.5Hz to a sample of saturated sand. Figure 5.17 shows cyclic torsion loading versus number of cycles to liquefaction in HCT test. In the first 4 cycles of Fig.5.17, the amplitude of cyclic loading was held constant without noticeable sample deformation with increasing number of cycles. It means soil still was strong. However, in 5 th cycles of Figure 5.17, the amplitude of cyclic loading dropped to 25Ibf ft. The amplitude of cyclic loading begun to decreased with increasing number of cycles after 4th cycles. After 4th cycles of Figure 5.17, the amplitude of cyclic loading rapidly dropped with increasing number of cycles. That means the both sample were liquef ied and too soft in the last few cycles. Figure 5.17 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test. (Jungang Liu, 2019 ) -50 -40 -30 -20 -10 0 10 20 30 40 50 0 2 4 6 8 10 12 Cyclic Torque (Ibf ft) Number of cycles to liquefaction Liquefaction happens

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178 Excess Pore Water Pressure versus Number of Cycles to liquefaction Figure 5.18 shows excess pore pressure versus number of cycles in Hollow Cylinder test. Excess pore water pressure starts to build up with the number of torsion loading cycles increase in the first 4 cycles. During the first 4 cycles of cyclic torsion appl ication, the sample was no noticeable deformation although the pore water pressure built up gradually. Nevertheless, the pore water pressure rapidly increased to the equal to the externally applied effective stress in the 5 th cycle. It showed that the soil had liquefied after the 5 th cycle. Figure 5.18 Pore water pressure change versus Number of cycles to liquefaction in CHCT. (Jungang Liu, 2019 ) 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore water pressure (psi) Number of cycles to liquefaction Liquefaction happens

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179 Cyclic Shear Stress versus Shear Strain The cyclic shear stress shear strain graph is also shown on Figure 5 .19. For the first 3 cycles the curves are close together, but as the sample approaches failure the strains increase and the hysteresis loops open up quickly. In the first 3 cycles, the ra nge of shear strain was 5.0 to +5.0. It means that the sample was no noticeable deformation. In last few cycles, the amplitude of deviator stress decreased with increasing the effective mean stress. It can be seen that the sample was softened and large fl ow deformation took place with increasing number of cycles. In 5th cycle, the loops begun flat shape, and also the amplitude of deviator stress rapidly dropped 30 percent. The sample developed large strains which, in the 5 th cycle, exceeded 50 percent duri ng the last three cycles. That means that the sample had liquefied. In last two cycles of Figure 5.18, the ranges of shear strain were 10% to + 10% . Figure 5.19 Cyclic shear stress versus shear strain in Hollow Cylinder Test. (Jungang Liu, 2019 ) -6 -4 -2 0 2 4 6 -15 -10 -5 0 5 10 15 Shear Stress (psi) Shear Strain (%) Liquefaction

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180 Stress Path The q mean effective stress failing and the amplitude of cyclic shear stress begun to drop. When the soil sample had liquefied, pore water pressure equaled to the external ly applied effective stress . In the figure ( ) versus q( q = shear stress q at 4.5 psi applied on the soil sample. In 5th cycle of figure 5.19, the amplitude of cyclic shear stress rapidly dropped 33 percent. At the s ame time, the effective mean stress kept decrease. It means that the soil sample turned softer and softer. Figure 5.20 p versus q stress path in Hollow Cylinder Test. (Jungang Liu, 2019 ) -6 -4 -2 0 2 4 6 0 2 4 6 8 10 12 14 16 q(psi) p' (psi)

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181 Effect of Fines Content on Liquefaction Resistance It shows in t he Chapter 7.

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182 CHAPTER VI COMPARISON OF CYCLIC TRIAXIAL AND HOLLOW CYLINDER TEST RESULTS Introduction A series of isotropically consolidated undrained cyclic triaxial (CTT) and cyclic hollow cylinder tests (CHCT) were conducted to determine and compare the behavior and liquefaction resistance of soil specimens. In this research, two soil specimens were inc luded for comparing the liquefaction resistance of soil in CTT and CHCT. One of soil specimens was uniform medium clean Monterey No. 0/30 sand, another is the mixture of a uniform medium Monterey No. 0/30 sand and Leyden clay from Golden Colorado. Eighte en cyclic triaxial and seventeen hollow cylinder tests were performed on the uniform medium clean Monterey No. 0/30 sand. All thirty five specimens were prepared at four different relative densities of 30,45,50,60 percent and tested at frequency 0.5 Hertz. Table 6.1 detailed this comparative test program with information including specimen densities, effective stres s, and cyclic stress ratio s, number of cycles to liquefaction for both CTT and CHCT. The mixture of a uniform medium Monterey No. 0/30 sand and Leyden clay with five different percentages of fine content (5%, 10%, 15%, 25% and 35%) and plasticity index 20 . Twenty cyclic triaxial and twenty cyclic hollow cylinder tests were performed on the mixture samples. All fourty samples were prepared at two d ifferent relative densities of 30 and 60 percent and tested at frequency 0.5 Hertz. Table 6.2 detailed comparative test program. All triaxial specimens were prepared to attain 2 inches in diameter and 4 inches in length and hollow cylinder samples to inner diameter of 8 inches, outer diameter of 10 inches, and height of 10 inches.

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183 Table 6.1: Result of co mparative study on clean sand ( L iu, et al, 2017 ) Stress Ratio*: Cyclic stress ratio (cyclic triaxial test) = Cyclic stress ratio (cyclic hollow cylinder test) = Cyclic Triaxial Test Cyclic Hollow Cylinder Test Relative Density (%) (psi) Cyclic Stress R atio * Frequency (Hz) Number of Cycles to liquefaction Relative Density (%) Effective Stress (psi) Cyclic Stress R atio * Frequency (Hz) Number of Cycles to liquefaction 30.3 15 0.15 0.5 16 30.16 15 0.25 0.5 5 30.24 15 0.25 0.5 9 29.87 15 0.4 0.5 3 30.35 15 0.4 0.5 5 29.64 30 0.2 0.5 30 28.73 30 0.15 0.5 60 30.35 30 0.3 0.5 16 29.81 30 0.25 0.5 35 30.28 30 0.4 0.5 10 30.3 30 0.4 0.5 16 45.32 15 0.2 0.5 28 50.36 15 0.15 0.5 100 45.41 15 0.3 0.5 15 49.47 15 0.25 0.5 40 45.28 15 0.4 0.5 8 50.38 15 0.4 0.5 14 44.82 30 0.2 0.5 160 50.35 30 0.15 0.5 1000 44.79 30 0.3 0.5 64 49.25 30 0.25 0.5 250 45.38 30 0.4 0.5 28 48.97 30 0.4 0.5 40 60.28 15 0.2 0.5 68 60.14 15 0.2 0.5 145 60.47 15 0.3 0.5 28 60.25 15 0.3 0.5 60 60.57 15 0.4 0.5 12 59.75 15 0.4 0.5 26 59.75 30 0.2 0.5 643 60.38 30 0.2 0.5 1500 58.88 30 0.3 0.5 181 60.41 30 0.3 0.5 368 60.54 30 0.4 0.5 53 60.18 30 0.4 0.5 102

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184 Table 6.2: Result of Comparative Study on mi xture samples (Jungang Liu, 2019 ) Cyclic stress ratio *: Cyclic stress ratio (cyclic triaxial test) = Cyclic stress ratio (cyclic hollow cylinder test) = Cyclic Triaxial Test Cyclic Hollow Cylinder Test Relative Density (%) Effective Stress (psi) Cyclic Stress R atio * Fine Content (%) Number of Cycles to liquefaction Relative Density (%) Effective Stress (psi) Cyclic Stress R atio * Fine Content (%) Number of Cycles to liquefaction 30.01 15 0.4 5 10 30.16 15 0.4 5 8 30.11 30 0.4 5 30 29.08 30 0.4 5 28 30.67 15 0.3 10 10 30.39 15 0.3 10 7 30.4 30 0.3 10 30 30.21 30 0.3 10 26 30.46 15 0.2 15 10 29.44 15 0.2 15 7 30.1 30 0.2 15 35 29.8 30 0.2 15 26 29.3 15 0.3 25 14 30.6 15 0.3 25 7 29.76 30 0.3 25 36 31.73 30 0.3 25 26 29.88 15 0.4 35 14 29.73 15 0.4 35 8 30.24 30 0.3 35 32 29.7 30 0.3 35 27 59.97 15 0.4 5 55 60.21 15 0.4 5 19 60.01 30 0.4 5 90 61.43 30 0.4 5 40 59.71 15 0.3 10 56 61.9 15 0.3 10 18 59.63 30 0.3 10 95 61.01 30 0.3 10 39 59.67 15 0.2 15 55 61.9 15 0.2 15 18 60.03 30 0.2 15 98 59.77 30 0.2 15 38 59.92 15 0.3 25 60 62.32 15 0.3 25 19 60.15 30 0.3 25 97 58.94 30 0.3 25 38 60.09 15 0.4 35 48 60.59 15 0.4 35 18 60.15 30 0.4 35 88 59.71 30 0.4 35 39

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185 Cyclic Strength Correction Factor between Cyclic Stress Ratio Causing Liquefaction in the Field and Cyclic Stress Ratio Causing Liquefaction of Triaxial Test Sample in the Laboratory Seed (1979) found that the cyclic stress ratio causing liquefaction under multidimensional shaking conditions in the field is related to the cyclic stress ratio ca using liquefaction of a triaxial test sample in the laboratory by expression: ( ) l ( ) l triaxial Where = 0.57 for K 0 = 0.4 = 0.9 to 1 for K 0 = 1 Table 6.3 Values of C r (some researchers) Source Equations C r for K 0 = 0.4 Cr for K 0 = 1.0 Finn, et al C r = (1+ K 0 )/2 0.7 1.0 Seed and Peacock Varies 0.55 0.72 1.0 Castro C r = (1+ K 0 )/(3 0.69 1.15 In this research, a series of cyclic hollow cylinder and cyclic triaxial tests were performed to establish the following relation between the cyclic liquefaction resistance stresses ratios of these two types of tests, the former closely simulates the field shear wave propagation during r , in the following equation: ( ) l ( ) l triaxial The results of this comparison on the clean sand are su mmarized Table 6.4. As may be seen, the C r values range from 0.46 to 0.63 at the same relative densities of 30 percent and 60 percent, which falls within the range of values that Seed assessed in 1979. For the samples with different fines content, Table 6.5 showed results of the comparison. The C r values range from

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186 0.32 to 0.93 at soil samples with different fines content. For 30 percentage of relative density in table 6.5, The C r values range from 0.5 to 0.93, but C r values at 60 percent of relative density range from 0.32 0.44. Table 6.6 showed the correction coefficient matrix of dependent variable Cr and independent variables. After run the correction coefficient analysis, the dependent variable is correction factor Cr . Three independent variables are relative density in cyclic triaxial test , D r (CTT), number of cycles to liquefaction in cyclic triaxial test No. cycles (CTT) and relative density in cyclic hollow cylinder test, D r (CHCT). A linear regression model was obtained as r (CTT) 0.002*No.Cycles (CTT) 0.030*D r (CHCT). The R 2 value of the regression equation is 0.84.

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187 Table 6.4: C r values on the clean sand produced in Laboratory by CTT and CHCT ( Liu, et al, 2017 ) Number of Cycles to liquefaction Effective S tress (psi) Cyclic Stress R atio Target Relative Density (%) C r Cyclic Triaxial Test Cyclic Hollow Cylinder Test Cyclic Triaxial Test Cyclic Hollow Cylinder Test Cyclic Triaxial Test Cyclic Hollow Cylinder Test 9 5 15 15 0.25 0.25 30 0.56 5 3 15 15 0.4 0.4 30 0.60 60 30 30 30 0.15 0.2 30 0.50 35 16 30 30 0.25 0.3 30 0.46 16 10 30 30 0.4 0.4 30 0.63 100 28 15 15 0.2 0.15 50 (45) 0.28 40 15 15 15 0.3 0.25 50 (45) 0.38 14 8 15 15 0.4 0.4 50 (45) 0.57 1000 160 30 30 0.2 0.2 50 (45) 0.16 250 64 30 30 0.3 0.3 50 (45) 0.26 40 28 30 30 0.4 0.4 50 (45) 0.70 145 68 15 15 0.2 0.2 60 0.47 60 28 15 15 0.3 0.3 60 0.47 26 12 15 15 0.4 0.4 60 0.46 1500 643 30 30 0.2 0.2 60 0.43 368 181 30 30 0.3 0.3 60 0.49 102 53 30 30 0.4 0.4 60 0.52

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188 Table 6.5: C r values on the mixture samples produced in laboratory by CTT and CHCT (Jungang Liu, 2019 ) Number of Cycles to liquefaction Effective S tress (psi) Cyclic Stress R atio Fine Content (%) Target Relative Density (%) Cyclic Triaxial Test Cyclic Hollow Cylinder Test Cyclic Triaxial Test Cyclic Hollow Cylinder Test Cyclic Triaxial Test Cyclic Hollow Cylinder Test Cyclic Triaxial Test Cyclic Hollow Cylinder Test 10 8 15 15 0.4 0.4 5 5 30 0.8 30 28 30 30 0.4 0.4 5 5 30 0.93 10 7 15 15 0.3 0.3 10 10 30 0.7 30 26 30 30 0.3 0.3 10 10 30 0.87 10 7 15 15 0.2 0.2 15 15 30 0.7 35 26 30 30 0.2 0.2 15 15 30 0.74 14 7 15 15 0.3 0.3 25 25 30 0.5 36 26 30 30 0.3 0.3 25 25 30 0.72 14 8 15 15 0.4 0.4 35 35 30 0.57 32 27 30 30 0.3 0.3 35 35 30 0.84 55 19 15 15 0.4 0.4 5 5 60 0.35 90 40 30 30 0.4 0.4 5 5 60 0.44 56 18 15 15 0.3 0.3 10 10 60 0.32 95 39 30 30 0.3 0.3 10 10 60 0.41 55 18 15 15 0.2 0.2 15 15 60 0.33 98 38 30 30 0.2 0.2 15 15 60 0.39 60 19 15 15 0.3 0.3 25 25 60 0.32 97 38 30 30 0.3 0.3 25 25 60 0.39 48 18 15 15 0.4 0.4 35 35 60 0.38 88 39 30 30 0.4 0.4 35 35 60 0.44

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189 (Jungang Liu, 2019 )

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190 Comparison on Both Test Results Cyclic Torque, Cyclic Axial Load versus Number of Cycles to Liquefaction In the Hollow Cylinder Test, the height of sample is 10.00 in, outside diameter is 10.00 in and inside diameter is 8.00in. The mass of sample is 6900.5g, and its relative density is 30.0%. effective stress is 80psi, its back pressure is 65 psi, its S.R. is 0.25. However, In Cyclic Triaxial Test, the height of sample is 4.0 inch; its diameter is is 0.96 when its effective stress is 80psi, and its back pressure is 65psi, its S.R. is 0.25. In HCT test, a cyclic torsion load of constant amplitude (40Ibf ft) was applied with a frequency of 0.5Hz to a sample of saturated sand. Figure 6.1 shows cyclic torsion loading versus number of cycles to liquefaction in HCT test. In Cyclic Triaxial test, a cyclic axial load of constant amplitude (24 Ib) was applied on top of soil specimen with a frequency of o.5 Hz. Figure 6.2 shows cyclic load versus number of cycles to liquefa ction in Cyclic Triaxial Test. In the first 8 cycles of Fig.6.2, the amplitude of cyclic loading was constant with increasing number of cycles because the sample was not noticeable deformation. It means soil still was strong. However, in 5 th cycles of Figu re 6.1, the amplitude of cyclic loading dropped to 25Ibf ft. The amplitude of cyclic loading begun to decreased with increasing number of cycles after 4th cycles. In cyclic triaxial test, more of numbers cycles to liquefaction than in HCT test. In cyclic t riaxial test, soil specimen stronger than in HCT test under the same cyclic stress ratio and relative density. After 4th cycles of Figure 6.1 and 9 th cycles of Figure 6.2, the amplitude of cyclic loading rapidly dropped with increasing number of cycles. T hat means the both sample were liquefied and too soft in the last few cycles.

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191 Figure 6.1 Cyclic Torque versus Number of cycles to liquefaction in Hollow Cylinder Test ( Liu , et al, 2017 ) -50 -40 -30 -20 -10 0 10 20 30 40 50 0 2 4 6 8 10 12 Cyclic Torque (Ib ft) Number of cycles to liquefaction Cyclic Torque vs Number of Cycles to liquefaction Liquefaction

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192 Figure 6.2 Cyclic Axial Load versus Number of cycles to liquefaction in Cyclic Triaxial Test. ( Liu , et al, 2017 ) liquefaction -28 -24 -20 -16 -12 -8 -4 0 4 8 12 16 20 24 28 0 2 4 6 8 10 12 14 16 18 20 Cyclic Axial Load(Ib) Number of cycles to liquefaction Cyclic Triaxial Test Dr = 30% Frequency = 0.5 Hz S.R.=0.25 Deviator Stress = 7.5 psi E.S.=15psi

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193 Pore Water Pressure versus Number of Cycles to Liquefaction Figure 6.3 shows excess pore pressure versus number of cycles in Hollow Cylinder test. During the first 4 cycles of cyclic torsion application, the sample showed no noticeable deformation although the pore water pressure built up gradually. However, during the 5th stress cycle, the pore pressure suddenly increased to a value equal to the externally applied effective stress . In fact, the soil had liquefied and the effective stress had been reduced to zero. Over a wide range of strains, the soil could be obse rved to be in a fluid condition. Pore water pressure continues to build up steadily as the number of stress cycles increase, until there is a sudden increase denoting the onset of initial liquefaction. The different values of pore water pressure developed during increases and decreases in deviator stress reflect the influence of the applied stress conditions. Figure 6.4 shows pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. Compare to hollow cylinder test, it nee d more number of stress cycles to make pore pressure increased to a value equal to the externally applied effective stress . In HCT test, pore water pressure of soil specimen more faster build up to be equal to effective stress 15 psi than in CTT test.

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194 Figure 6.3 Pore water pressure change versus Number of cycles to liquefaction in HC Test. ( Liu , et al, 2017 ) 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 u(psi) Number of cycles to liquefaction Pore water pressure change vs Number of cycles to liquefaction Liquefaction

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195 Figure 6.4 Pore water pressure change versus Number of cycles to liquefaction in Cyclic Triaxial Test. ( Liu , et al, 2017 ) Liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 pore pressure (psi) Number of cycles to liquefaction Test #4 Dr=30% Frequency = 0.5 Hz S.R.=0.25 Deviator Stress = 7.5psi C.P.=15psi

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196 Cyclic Shear Stress, Cyclic Deviator Stress versus Shear Strain, Axial Strain The shear stress shear strain graph is also shown on Figure 6. 5 . For the first 3 cycles the curves are close together, but as the sample approaches failure the strains increase and the hysteresis loops open up quickly. In the first 3 cycles, the range of shear strain was 5.0 to +5.0. It means that the sample was no n oticeable deformation. In last few cycles, the amplitude of deviator stress decreased with increasing the effective mean stress. It can be seen that the sample was softened and large flow deformation took place with increasing number of cycles. In 4 th cycl e, the loops begun flat shape, and also the amplitude of deviator stress rapidly dropped 30 percent. The sample developed large strains which, in the 5 th cycle, exceeded 50 percent during the last three cycles. That means that the sample had liquefied. In last two cycles of Figure 6.5, the ranges of shear strain were 10% to + 10% . In HCT test, a cyclic shear stress of constant amplitude ( 4.5 psi ) was applied with a frequency of 0.5Hz to a sample of saturated sand . However, a cyclic deviator stress (7. 5 psi) was applied on the top of soil specimen with the same frequency of 0.5 Hz. I n CTT, the range of axial strain was 0. 15 to +0. 15 in the first three cycles. It is not noticeable deformation like in hollow cylinder test . In last few cycles of Figure 6. 6, the amplitude of deviator stress decreased with increasing the effective stress , and also soil sample was softened. In last cycles of Figure 6.6, the sample got larger strain, which was 0.3 to 0.3. It means that the sample had liquefied.

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197 Figure 6.5 Cyclic shear stress versus shear strain in Hollow Cylinder Test ( Liu , et al, 2017 ) -6 -4 -2 0 2 4 6 -15 -10 -5 0 5 10 15 Shear Stress (psi) Shear Strain (%) Cyclic Shear Stress vs Shear strain Liquefa ction

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198 Figure 6.6 Cyclic deviator stress versus axial strain in Cyclic Triaxial Test. ( Liu , et al, 2017 ) -10 -8 -6 -4 -2 0 2 4 6 8 10 -0.5 -0.4 -0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4 0.5 (psi) a Liquefaction

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199 Stress Path Q versus stress (q) started at 4.5psi applied on the sample in the beginning of Hollow Cylinder test. In the first 4 cycles of shear stress application, the amplitude of cyclic shear stress kept constant with decreasing the effective mean stress. However, at the beginning of cyclic triaxial test, the also deviator stress (7.5psi) applied on the sample, and during first 8 cycl es of stress application, the amplitude of cyclic deviator stress kept constant with decreasing the effective mean stress. The q mean effective stress iquefaction condition at which time the sample starts failing and the amplitude of cyclic shear stress begun to drop. When the soil sample had liquefied, pore water pressure equaled to the externally applied effective stress . In 5th cycle of hollow cylinde r test, the amplitude of cyclic shear stress rapidly dropped 33 percent. However, in the cyclic triaxial test of figure 6.8, the amplitude of cyclic deviator stress dropped 10 percent after 8 th cycle. It means pore water pressure increased closed to the value of applied effective stress . After 5th cycle of hollow cylinder test (8 th cycle of cyclic triaxial test), the amplitude of cyclic shear stress continued to drop, at the same time, the effec tive mean stress kept decreased. It means that the sample turned softer and softer.

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200 Figure 6.7 p versus q stress path in Hollow Cylinder Test. ( Liu , et al, 2017 ) -6 -4 -2 0 2 4 6 0 2 4 6 8 10 12 14 16 q(psi) p' (psi) Liquefaction

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201 Figure 6.8 p versus q stress path in Cyclic Traixial Test. ( Liu , et al, 2017 ) -10 -8 -6 -4 -2 0 2 4 6 8 10 0 2 4 6 8 10 12 14 16 18 20 q (psi) p' (psi) liquefaction

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202 CHAPTER VII THRESHOLD FINES CONTENT Previous Studies In Dr. Hsing Denver sand at a constant overall void ratio does not necessarily cause increase in liquefaction resistance, and also at a constant overall void ratio, the increase in fines content at a constant plasticity index up to 30% in the clean sand cause a decrease in li quefaction resistance, beyond 30%, the further increase in fines content results increase in liquefaction resistance. Based on run triaxial test with samples tested under15 psi confining pressure in the medium sand with 5% fines and plasticity indices ran ging from 0 to 40, Nien Yin Chang (1990) showed that the trend of increasing liquefaction resistance with increasing plasticity index is more obvious, and also comparison between the effect of fine contents and the effect of plasticity indices of fines on the liquefaction resistance of soil indicated that the effect of fine contents is more significant than the effect of plasticity indices. I.M.Idriss and R.W.Boulanger (2004) showed that cyclic stress ratio (CSR) increased with increased fine content perc ent under the same modified standard penetration blows count. case history points for cohesionless soils with 5 %< FC <15%, while Figure 7.3 showed the cases fo However, Chang (1990) showed the clean medium sand has the strongest liquefaction resistance and as the fine content increases, the liquefaction resistance decreases until the fine content reaches approximately 26%, as indicated by the l iquefaction potential curves shifting

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203 toward the left. Then the trend reverses itself and soils begin to gain strength as the fine content further increases, as indicated by the curves shifting toward the right. In Figure 7.4, it showed for the medium sand test series, exactly the similar trend of decreasing resistance and then increasing resistance with the increasing fine content. Each curve in Figure 7.4 gave the stress ratios required to achieve initial liquefaction in ten cycles of loading for soils wi th the same plasticity index for each curve. The collection of curves also indic ates the existence of a threshold fine content, below which, the resistance decrease, and, above which, the resistance increases with increasin g fine content, and the threshold fine content decrease with increasing plasticity of soils. Figure 7.4, 7.5, and 7.6 se em to indicate that the threshold fine content increases as the number of cycles required reaching initial liquefaction increases. Tzuo Shin Ueng, Chia Wen Sun and C hieh When Chen (2013) showed that as the fine content increased, cyclic resistance ratio decreased until the fine content approximately 20% in the Figure 7.7. And also indicated that the effect of fines on the liquefaction resistance of a soil was more pro minent using (FC) 400 (passing the No.400 sieve) than (FC) 200 (passing the No. 200 sieve) , probably due partly to the different in plasticity of the fines. Yong Wang and Yanli Wang (2010) showed that with a constant dry density, the liquefaction resis tance first increased and then decreased with the increase of the fines content, when the relative density reached the minimum value at the fines content of 30% and the liquefaction resistance also reached the minimum value at the same fines content in Fig ure 7.8. Mehmet Murat Monkul and Jerry A. Yamamuro (2011) showed that relative density alone cannot be a consistent comparison basis for the influence of fines content on liquefaction potential of sand in Figure 7.9, and also the D 50 sand /d 50 silt ratio becomes larger (for SilCoSil and

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204 Potsdam fines), the void ratio decreases consistently with increasing fines content, show as Figure 7.10. Polito and Martin (2001) used Monterey and Yatesville sand with Yatesville silt to study the effects of non plastic fines on the liquefaction resistance of sands. Their study showed the transitional fines content around 35 %, based on the relationship between cyclic stress ratio and void ratio. Thevanayagam et al. (2002) used foundry sand and silica fines to study the u ndrained fragility of clean sands, silty sands, and sandy silts. The transitional fines content based on the steady state line was around 40 %. Shaoli Yang, Suzanne Lacasse, and Rolf Sandven (2005) showed the curves, based on the cyclic test results, appear change direction at fines content of 30% and the curves for threshold fine content of 30% always give low the void ratio and low cyclic stress ratio. Chang (1990) showed that the medium sand with a small fine content of 5% reflects in the irregular relationship between the stress ratio required to reach initial liquefaction versus plasticity index, and also comparison between the effect of fine contents and the effect of plasticity indic es of fines on the liquefaction resistance of soils indicated that the effect of fine contents is very much more significant than the effect of plasticity indices.

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205 Fig. 7.1 SPT case histories of cohesionless soils with FC 35% and the NCEER Workshop (1997) curve and the recommended curves for both clean sand and for FC = 35% for M = 7½ and ' vo = 1 atm ( I.M.Idriss and R.W.Boulanger, 2004).

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206 Fig.7.2 SPT case histories of cohesionless soils with 5%
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207 Fig.7.3 : SPT case histories of cohesionless soils with 15%
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208 Figure 7.4 St ress Ratio Required to reach Liquefaction in 10 Cycles versus Fine Content Under Confining pressure 15 psi ( N.Y. Chang, 1990) Figure 7.5 Stress Ratio Required to reach Liquefaction in 30 Cycles versus Fine Content Under Confining pressure 15 psi ( N.Y.Ch ang, 1990)

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209 Figure 7.6 Stress Ratio Required to reach Liquefaction in 100 Cycles versus Fine Content Under Confining pressure 15 psi ( N.Y.Chang, 1990) Figure 7.7 Cyclic resistance ratio versus equivalent (FC) 400 (Tzuo Shin Ueng, Chia Wen Sun a nd Chieh When Chen, 2013)

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210 Fig.7.8 Variation in Liquefaction Resistance with Fines Content for N l =20 ( Yong Wang and Yanli Wang, 2010) Fig.7.9 Change of relative density and liquefaction potential with different fines contents and Silts for tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011)

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211 Fig.7.10 Change of void ratio and liquefaction potential with differen t fines contents and silts for tested specimens. (Mehmet Murat Monkul and Jerry A. Yamamuro, 2011)

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212 Threshold Fines Content Definition Research has shown that the behavior of the mixtures can be characterized into two groups, one is sand dominated, and the other is fines dominated. The intergrain state concept was proposed by Thevanayagam (1998). The intergrain void ratio (referring to the intergranular and interfine void ratio s) is used to indicate the state of the sand and silt mixtures instead of the global void ratio. For sand with low fines content, the intergranular void ratio e c is defined as: e c = where fc is fines content (in decimal) and e is the void ra tio of the silty sand. For sand with high fines content, the interfine void ratio e f is defined as: e f = Threshold fines content is a transitional fines content, sand dominated behavior passes to fines dominated behavior when the fines content is b eyond the threshold fines content. Threshold fines content (FC th ) is expected to occur when the shear response of the mixtures is expected to weaken, fine grains plays a primary role. FC th % = % where e max,HF = the maximum void ratio of the pure silt above which it has no appreciable strength. Factors Effects Threshold Fines Content Factors from Soil Properties The threshold fines content depend on many factors from soil properties such as the fine part icles, volumes of solids and void, the values of e max and e min , global void radio (e) and the characteristics of fines and coarse grains Fines p articles In the intergrain state concept, the fine particles are regarded as void when the fines content is low (Fig. 7.11), and the fines are assumed to not participate in the resistance of shearing. When the fines content is high, the sand grains are regarded as void, and the sand particles are assumed to not contribute to the shearing resistance (Fig. 7.11).

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213 Fig. 7.11 (a) Sand with low fines content (b) Sand with high fines content (Shaoli Yang, Suzanne Lacasse and Rolf Sandven, 2005) Volume solids and v oid When mixing spherical particles of two different sizes, the packing will be affected by the proportion of large size and small size spheres in the total volume of solids as well as by the relative size of the large and small spheres. Misko Cubrinovski and Kenji Ishihara (2002) showed how the volumes of solids and v oid vary with a change in the percentage of the small size particles in Figure 7.12. In Figure 7.12, point L means the densest possible packing of the larger spheres. In the beginning, adding smaller size particles into the densest packing of large spheres cause to a decrease in the volume of voids because the small spheres fill in the voids among the larger particles. This filling of voids phase is showed in the diagram with the path L T. Upon adding small particles beyond a certain percentage correspondin g to point T, a reverse trend is observed in which the volume of voids increase with the percentage of the s mall size fraction. In this so called replacement of solids phase, the large size particles are pushed apart and gradually replaced by the small siz e spheres until the entire volume of solids are comprised of smaller particles at point S. It is indicated that e min decreases in the course of the filling of voids process and reaches its minimum value of e min(T) at the threshold percentage corresponding to T, and then e min steadily increases during the replacement of solids process at the path T S in the Figure7.12(b). Misko Cubrinovski and Kenji Ishihara (2002) exposed that fines containing sands are that the threshold percentage of fines at which the fi lling process is reversed into the replacement process is significantly smaller than 50%.

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214 Figure 7.12 Effects of fines on Binary Packing of Spherical Particles: (a) variation in the volume of voids and solids (b) variation on e min . (Misko Cubrinovski and Kenji Ishihara, 2002) e max and e min Misko Cubrinovski and Kenji Ishihara (2002) showed that the values of e max and e min , have been determined by the adopted laboratory procedures, as a function of the fines contents for each of the composite soils, as shown in Figure 7.13. In Figure 7.13, it is observed that e max and e min decreased as the fines content increased from 0% t o 20%. However, e max and e min reached the lowest void ratio within the range of 20 to 40% because soil had a transition from the filling of voids to the replacement of solids process. After 40% fines, e max and e min are seen to steadily increase until they reach the highest values at 100% fines content.

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215 Figure 7.13 Variation in e max and e min with fine content of mixtures of Cambria Sand and Nevada fines. (Misko Cubrinovski and Kenji Ishihara, 2002) Global v oid r adio, c haracteristics of fines and coarse g rains S. Thevanayagam; T. Shenthan; S. Mohan and J. Liang (2002) found that the value of threshold fines content (FC th ) depends on global void radio (e) and the characteristics of fines and coarse grains. It was found that at the same e and confining stress, the collapse potential (fragility) of silty sand increases with an increase in fines content (FC) due to a reduction in intergranular contact between the coarse grains. At FCFC th ), fine grain friction plays a primary role and dispersed coarse grains provide a beneficial, secondary reinforcement effect. They showed that the microstructure of a granular mix, which can be constituted in many different wa ys with different types of intergrain contacts, leads to different undrained shear responses in the Figure 7.14. There are three extreme limiting categories of microstructure: (a) primarily the coarse grains are in contact [cases (i) (iii) in Fig. 7.14 (a )], (b) primarily the fine grains are in contact with each other [case (iv) in Fig. 7.14 (b)], (c) a layered system [Fig. 7.14 (c)].

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216 Figure 7.14 Intergranular soil mix classification. (S. Thevanayagam; T. Shenthan; S. Mohan and J. Liang, 2002) Factors from Laboratory Tests There are many factors a ffect ing threshold fines content in the laboratory test. They includes relative density, cyclic stress ratio, consolidation pressure and number of cycles to liquefaction. Based on all laboratory test results from cyclic triaxial tests and cyclic hollow cylinder te sts, they indicated that threshold fines content was happened and the value was 15 % . Table 7.1 showed threshold fines content, test parameters and soil properties in cyclic traixial test and hollow cylinder test. Table 7.2 showed the correction coeffici ent matrix of dependent variable threshold fines content and independent variables. After run the correction coefficient analysis, the dependent variable is threshold fines content, TFC. Three independent variables are relative density, D r , number of cycle s to liquefaction, No. and consolidation pressure, CP. A linear regression model was obtained as follows: TFC = 14.88 0.022*D r +0.027*No. 0.009*CP. The R 2 value of the regression equation is 0.44.

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217 Table 7.1 Threshold fines content in cyclic triaxial test and cyclic h ollow cylinder t est (Jungang Liu, 201 9 ) Threshold Fines Content (%) Relative Density (%) No. Cycles to Liquefaction Consolidation Pressure (psi) Cyclic Stress Ratio Cyclic Traixial Test 14 30 10 15 0.2 16 30 35 30 0.2 15 45 35 15 0.2 15 45 50 30 0.2 14.5 60 50 15 0.2 15.5 60 90 30 0.2 Cyclic Hollow Cylinder Test 14 30 8 15 0.2 14 30 27 30 0.2 14.5 60 20 15 0.2 14 60 40 30 0.2 Table 7. 2 Correlation coefficient matrix of threshold fines content and laboratory factors (Jungang Liu, 201 9 ) Results of Laboratory Test All Soil S amples Results in Cyclic Triaxial Test A series of isotropically consolidated undrain cyclic triaxial tests were conducted to investigate the effect of fine contents on the liquefaction resistance of soils. Ninety six cyclic

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218 triaxial tests were performed on the uniform Monterey No. 0/30 sand with six different percenta ges of fine content (5%, 10%, 15%, 25%, 35% and 45%) and plasticity index 20. Stress Ratio for 10, 35, 50 and 90 Cycles to Liquefaction vs Different Percent of Fine Content The figures of stress ratio vs number of cycles to liquefaction for soil with dif ferent percent of fine contents to determine the stress ratio required to reach liquefaction in 10, 35, 50 and 90 cycles for all sample tested. In the Figure 7.15, it shows the relationship between the stress ratio for 10 and 35 cycles to liquefaction and soil with different percent of fine contents (5%, 10%, 15%, 25%, 35% and 45%) and the same PI. In the Figure 7.15, soil samples prepared at relative density of 30% and run under 15 psi, 30psi of effective consolidation stress. The Curve of Figure 7.15 indi cated that the liquefaction resistance decreases as the fine content increase until the fine content reaches 15%. The soil liquefaction resistance has the lowest value at 15 % of fine content. After then the trend reverse itself and soil begin to gain stre ngth with fine content increase. Figure 7.16 present that the relationship between the stress ratios required reaching liquefaction in 35 and 50 cycles and soil with various percent of fines content (5%, 10%, 15%, 25% and 35%). In the Figure 7.16, all sam ples prepared at relative density of 45% and run under 15 psi, 30psi of effective consolidation stress. The curves in figure 7.15 and 7.16 are very similar sharp. Figure 7.17 present that the relationship between the stress ratios required reaching liqu efaction in 50 and 90 cycles and soil with various percent of fines content (5%, 10%, 15%, 25% and 35%). In the Figure 7.17, all samples prepared at relative density of 60% and run under 15 psi, 30psi of effective consolidation stress. Figure 18 gave the same curve sharp as in the Figure 7.15 and 7.16.

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219 Figure 7.15 Stress ratio require for liquefaction in 10 and 35 cycles versus variable fine content (M L mixing samples at relative density 30% and confining pressure 15,30 psi) in cyclic triaxial tests. (Jungang Liu, 2019 ) Figure 7.16 Stress ratio require for liquefaction in 35 and 50 cycles versus variable fine content (M L mixing samples at relative density 45% and confining pressure 15,30 psi) in cyclic triaxial tests. (Jungang Liu, 2019 ) 0 0.1 0.2 0.3 0.4 0.5 0 5 10 15 20 25 30 35 40 45 50 Stress Ratio Fines Content (%) Dr=30% PI=20 E.S.=15psi No.cycles =10 Dr=30% PI=20 E.S.=30psi No.cycles=35 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 5 10 15 20 25 30 35 40 45 50 Stress Ratio Fine Content (%) Dr=45% ES=15psi PI=20 No.Cycles=35 Dr=45% E.S.=30psi PI=20 No.Cycles=50

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220 Figure 7.17 Stress ratio require for lique faction in 50 and 90 cycles versus variable fine content (M L mixing samples at relative density 60% and confining pressure 15,30 psi) in cyclic triaxial tests. (Jungang Liu, 2019 ) All S oil S amples Results in Cyclic Hollow Cylinder Test Twenty hollow cylinder tests were performed to investigate the effect of fine contents on the liquefaction resistance of soil. All samples are the mixture of uniform Monterey No. 0/30 sand with five different percentages of fine content (5%, 10%, 15%, 25% and 35%) and plasticity index 20. All hollow cylinder specimens were prepared at two different relative densities of 30, 60 percent and run under 15 psi, 30 psi of effective consolidation stress. Figure 7.18 present that the relationship between the stre ss ratios required reaching liquefaction in 8 cycles and soil with various percent of fines content (5%, 10%, 15%, 25% and 35%). In the Figure 7.18, all samples prepared at relative density of 30% and run under 15 psi, 30psi of effective consolidation stre ss. In the cyclic hollow cylinder test, the curve of figure 19 showed the same relationship between stress ratio and fines content as in the cyclic triaxial test. The liquefaction resistance began to decrease with fine content increase until fine content reaches 15%, after that soil got stronger as the fine content increases. 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 5 10 15 20 25 30 35 40 Stress Ratio Fine Content (%) Dr=60% E.S.=15psi PI=20 No.Cycles=50 Dr=60% E.S.=30psi PI=20 No.Cycles=90

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221 In the Figure 7.19, 7.19, 7.20 and 7.21, they showed that the same results as in the cyclic triaxial test, and proved that the threshold fines content is expected to occur. In this r esearch, it showed that the threshold fines content is 15% on cyclic triaxial and cyclic hollow cylinder tests. Figure 7.18 Stress ratio require for liquefaction in 8 cycles versus variable fine content (mixing samples at relative density 30% and conf ining pressure 15psi) in cyclic hollow cylinder tests. (Jungang Liu, 2019 ) 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 5 10 15 20 25 30 35 40 Stress Ratio Fine Content(%)

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222 Figure 7.19 Stress ratio require for liquefaction in 27 cycles versus variable fine content (mixing samples at relative density 30% and confining pressure 30 psi) in cyclic holl ow cylinder tests. (Jungang Liu, 2019 ) 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 5 10 15 20 25 30 35 40 Stress Ratio Fine Content(%)

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223 Figure 7.20 Stress ratio require for liquefaction in 20 cycles versus variable fine content (mixing samples at relative density 60 % and confining pressure 15 psi) in cyclic hollow cylinder tests. (Jungang Liu, 2019 ) 0 0.1 0.2 0.3 0.4 0.5 0 5 10 15 20 25 30 35 40 Stress Ratio Fines Content (%) CHCT No.Cycles =20 Dr=60% Consolidation stress=15psi

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224 Figure 7.21 Stress ratio require for liquefaction in 40 cycles versus variable fine content (mixing samples at relative density 60 % and confining pressure 30 psi) in cyclic hollow cylinder tests. (Jungang Liu, 2019 ) 0 0.1 0.2 0.3 0.4 0.5 0 5 10 15 20 25 30 35 40 Stress Ratio Fines Content (%) CHCT No.Cycles =40 Dr=60% Consolidation Stress=30 psi

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225 CHAPTER VIII EXCESS PORE PRESSURE GENERATION Excess Pore Pressure Generation from Laboratory Test R esults Previous Studies In the 1970s, Seed et al.(1975b) developed an empirical model for predicting r u using data from tests performed on clean sands. In their model, r u is a funct ion of the cycle ratio, which is the ratio of the number of applied uniform cycles of loading ( N ) to the number of cycles required to cause liquefaction in the soil ( N l r u = arcsin (2*( 1) Accord ing to Polito et al. (2008), the value of is a function of fine content (FC), relative density ( D r), and CSR, and cannot be assumed equal to 0.7 for all cases. Booker et al. (1976) proposed an alternative, somewhat simplified version of this equation r u = arcsin ( Each of the above equations makes use of two calibration parameters (i.e., N l can be determined from stress controlled cyclic triaxial tests, as well as other types of undrained cyclic tests. Nonlinear mixed effect (N LME) models were used in regression analyses to develop geotechnical/ earthquake data include Abrahamson and Silva (1996), Liu et al.(2001), and Rathje et al. (2004). Several forms of equations were used in the regression analyses, with the following tot ): 1 *FC+c 2 *D r +c 3 *CSR+c 4 where Dr=relative density in percent; CSR=cyclic stress ratio; FC=fines content in percent; an d c 1 , c 2 , c 3 , c 4 , are regression coefficients (for FC<35%: c 1 =0.01166; c 2 =0.007397; c 3 =0.01034; and c 4 =0.5058; and for FC 35%: c 1 =0.002149; c 2 c 3 =1.667; and c 4 =0.4285) Green et al. (2000) developed the the Green Mitchell Polito model (GMP) model, which is an empirical expression that relates r u to the energy dissipated per unit volume of soil (i.e., unit energy). The GMP model is a special case of the more general energy based model proposed by

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226 Berrill and Davis (1985). The GMP model was developed using data from tests performed on nonplastic silt sand mixtures that ranged in fines contents from clean sands to pure silts. The GMP model is r u = where W s =energy dissipated per unit volume of soil divided by the initi al effective calibration parameter. For general loadings, increments in W s can be related to stress conditions and increments in strain equation is dW s = ( d + d d + d where dW s =incremental dissipated energy normalized by the initial effective mean stress; =effective vertical stress; d =incremental vertical strain; =effective horizontal stress; d =incremental radia l strain; =horizontal shear stress acting on a plane having a vertical normal vector; d =incremental shear strain resulting from ; =vertical shear stress acting on a plane having a horizontal normal vector; d =incremental shear strai n resulting from ; and =initial effective stress. For undrained cyclic triaxial test loadings, W s can be computed numerically W s = )( where n =number of load increments to liquefaction; and =applied deviator stress at load increment i and i +1, respectively; and and =axial strain at load increment i and i +1, respectively. The pseudoenergy capacity (PEC) is determined from cyclic test data b y plotting r u versus the square root of W s . The square root of PEC is the value on the horizontal axis corresponding to the intersection of a straight line drawn through the origin and the point of ru =0.65 and a horizontal line drawn at r u =1.0. = where Dr=relative density in percent; and c 1 , c 2 , c 3 , and c 4 are regression coefficients ( c 1 c 2 =0.312; c 3 =0.0139; and c 4 tot ln(PEC)= 0.6591.

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227 For FC< 35%, PEC increases as Dr increases, and decreases as FC increases. In contrast, es and decreasing slightly as FC increases. Vucetic and Dobry (1986) developed a unique relationship among PWP ratio (ru), cyclic controlled cyclic triaxial com pression tests, as: r u,N = where r u,N = residual excess PWP ratio at cycle N; f = 1 or 2 depending on whether cyclic loading is generated by one or two directional loading; p, F, and s = curve fitting constants; and tvp = volumetric threshold shear strain, defined as the shear strain thresho ld below which no significant PWP is generated during cyclic loading. This shear strain falls between 0.01 and 0.02% for most sands (Dobry et al. 1982). Mei et al. (2015) recommended values of p = s =1 for the curve fitting parameters for clean sands. Figu re 8.1 presents their proposed correlation between the parameter F and soil index properties Dr (relative density) and CU (coefficient of uniformity). Figure 8.1 . Correlation to estimate parameter F in Vucetic Dobry PWP generation model . ( Vucetic and Dob ry, 1986) Martin, Finn and Seed, (1975) based on the compatibility of volume change of soil skeleton and the pore water developed a relation for the pore water pressure increment for each

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228 cycle of shear stress. For saturated sand under an undrained cyclic shearing loading, the total volumetric strain is the sum of the volumetric strains of solid particle and pore water. vd = the increase in residual pore pressure for the cycle; k w = bulk modulus of water; n e = porosity of sample; and r = tangent modulus of the one dimensional unloading curve at a point corresponding to the initial vertical effective stress, then consideri ng the unit volume of sand; Change of volume of voids = ; reduction in volume of sand structure due to vd ; and increase in volume of sand structure due to recoverable volumetric vr = , For saturated sample , K w = 4 10 7 psf, where r is generally of the order of 10 6 psf. Considering the relative orders of magnitude of the moduli, the water may be assumed to be effectively incompressible, and thus, under conditions of zero volume change: vd vr = Or r vr The key to the practical application of th vd have been found to be independent of vertical stress. Therefore, the theory in its simplest form implies vo has a recovera vro , then liquefaction will occur under an applied cyclic strain vd vr under drained conditions. Ueng, Wu, Lin and Yu (2000) showed that the change of effective stress, or pore water pressure of a saturated sand under the undrained condition, for a given shear strain increment becomes, = = u = [ ] d = effective normal stress on the shearing plane = basic friction angle between sand grains b = a positive value related to shearing plan orientation, friction angle, and sand fabric , n = porosity of the sand

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229 k w = bulk modulus of pore water k r = rebound modulus of volume change of the sand structure Chung, Yokel and Wechsler (1984) conducted three torsional resonant column tests on hollow cylindrical specimens. The specimens were prepared to 60% relative density and tested under a 96 kpa confining pressure. The correlation between the volumetric strain, as measured by pore water displacement, and the excess pore water pressure developed is shown in Figure 8.2. Strain controll ed cyclic triaxial tests reported by National Research Council (NRC,1985) in Figure 8.3. In the figure 8.3, it indicated that little pore pressure is generated for 10 cycles of t = 0.005% gives an upper boun d to this data. Based t = 0.005% was selected for calibration with the results of load controlled tests. Figure 8.2 . Excess Pore Water Pressure ratio vs. Volumetric Strain (Chung, R.M., Yokel, F.Y. and Wechsler, H. 1984)

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230 Figure 8.3 Excess pore water pressure build up Test data from NRC (1985) Excess P ore P ressure Generation from L aboratory T est R es ults A series of isotropically consolidated undrained cyclic triaxial (CTT) and cyclic hollow cylinder tests (CHCT) were conducted to determine excess pore water generation. All samples were prepared at three relative densities of 30%, 45% and 60%, mixed six different perc entages of fines contents (5%, 10%, 15%, 25%, 35% and 45%), and tested at different stress ratios and effective consolidation stress. Excess pore pressure ratio, R u , defined as the radio of excess pore water pressure, u, to initial effective 3 . The normalized excess pore water pressure ratio was expressed as and the normalized number of cyclic stress cycle was designated as . In the figure 8.4 and 8.5, they showed that the relationship between the normalized excess pore pressure ratio and the normalized number of cyclic stress cycle in cyclic triaxial tests and cyclic hollow cylinder tests. In the figure 8.6, it indicated that average excess pore pressure versus normalized number of cyclic stress cycles in all samples of cyclic triaxial tests and cyclic hollow cylinder tests.

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231 Figure 8.4 . The normalized excess pore water pressure ratio versus the normalized number of cyclic stress cycle with different percent of fines content and relative densities in cyclic triaxial test. a) relative density at 30% and different of fines content (0%, 5%,10%,15%,25%,35% and 45%). (Jungang Liu, 2019 ) 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Dr = 30% in CTT FC0% FC5% FC10% FC15% FC25% FC35% FC45%

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232 b) relative density at 45% and different of fines content (0%, 5%,10%,15%,25% and 35%). (Jungang Liu, 2019 ) c) relative de nsity at 60% and different of fines content (0%, 5%,10 %,15%,25% and 35%).(Jungang Liu, 2019 ) 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Dr = 45% in CTT FC0% FC5% FC10% FC15% FC25% FC35% 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Dr = 60% in CTT FC0% FC5% FC10% FC15% FC25% FC35%

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233 Figure 8.5 . The normalized excess pore water pressure ratio versus the normalized number of cyclic stress cycle with different percent of fines content and relat ive densities in cyclic hollow cylinder test. a) relative density at 30% and different of fines content (0%, 5%,10%,15%,25% and 35%).(Jungang Liu, 2019 ) 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Dr = 30% in CHCT FC0% FC5% FC10% FC15% FC25% FC35%

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234 b) relative density at 60% and different of fines content (0%, 5%,10%,15%,25% and 35%). (Jungang L iu, 2019 ) 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Dr = 60% in CHCT FC0% FC5% FC10% FC15% FC25% FC35%

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235 Figure 8.6 . Average excess pore pressure versus normalized number of cyclic stress cycles in all samples a) in cyclic triaxial test with different of fines content (0%, 5%, 10%, 15%, 25%, 35% and 45%).(Jungang Liu, 2019 ) 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Average Excess Pore Pressure for all samples in CTT FC 0% FC5% FC10% FC15% FC25% FC35% FC45%

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236 b) in cyclic hollow c ylinder test with different of fines content (0%, 5%,10%,15%,25% and 35%). (Jungang Liu, 2019 )8.2 Constitutive models for simulating pore pressure generation Constitutive Models for Simulating Pore Pressure G erneration Constitutive Model UBC3D PLM UBC2D is used for the UBCSAND model, has defined at first by Puebla et al. and then has used in FLAC software by Beaty and Byrne (Puebla et al., 1997, Beaty & Byrne, 1998). The UBCSAND model is a simple 2D model developed specially for estimate of liquefaction behavior of sand. Also the model has been verified in various applications related to liquefaction. The original 2D model uses a Mohr Coulomb yield function and a corresponding non associated plastic potential function. The flow rule is bas ed on the well dilatancy formulation with a modification (Rowe, 1962). UBCPLM model is based on UBCSAND model which has presented by Anteneh Biru Tsegaye (A.B. Tsegaye, 2010). UBCPLM model, primarily based on the elastoplastic functions mentioned accordingly far a generalized 3D formulation has been considered. The new model 0 0.2 0.4 0.6 0.8 1 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 u/ 3' N/Nliq Average Excess Pore Pressure for all samples in CHCT FC0% FC5% FC10% FC15% FC25% FC35%

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237 uses the Mohr Coulomb yield condition in a generalized stress space. The use of non associated plastic potential based on the same function as the yield function (w ith mobilized friction angle replaced by mobilized dilatancy angle) has been found to introduce non coaxially between the stress and the strain in the deviatoric plane. In 2013, Alexanderos Petalas and Vahid Galavi have introduced UBC3D PLM code for using in Plaxis (A, Petalas & V, Galavi, 2013). The UBC3D PLM combination has three aspects: a) UBC model; b) 3 Dimension; c) Past Liquefaction Model. The undrained behaviour of the soil is treated implicitly by the UBC3D PLM constitutive model. Therefore, the increment of the pore water pressure is computed at each step of the analysis. Considering a saturated soil specimen, the increments in total stress during loading is given by the following equation: v where Ku is the bulk modulus of the undrai v the volumetric strain of the soil as a whole. The effective stress increment can be computed as follows: v v its volumetric strain. The increments of the pore water pressur e is computed with the following equation: dp w = v where K w v is the volumetric strain of the fluid. The relationship between the total stresses, the effective stresses and the pore pr essure is undrained conditions requires that the equivalent fluid volumetric strain must be equal to the volumetric strain of the soil skeleton. w = K u This value is close to the upper limit (of 0.5) as water is almost incompressible. Using a value of 0.5 is to be avoided as this is known to cause numerical insta ratio the bulk modulus of the undrained soil is computed as follows:

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238 K u = Where G e is the elastic shear modulus. = In the latest version of the UBC3D the bulk modulus of water is dependent with the degree of saturation of the soil. Finn Constitutive model FLAC software (Fast Lagrangian Analysis of Continua) is a Finite Difference Method based program (FDM). According to FLAC guidance manual, there are several constitutive models that facilitate soil behavior under static and dynamic loadings (Itasca FLAC ma nual, 2008). Calculation of excess pore water pressure in the soil mass due to dynamic loading is the main factor in the modeling process of liquefaction phenomenon. FLAC has a constitutive model named Finn model which equations represented by Martin et a l. (1975) and Byrne (1991) into the standard Mohr Coulomb plasticity model. Using this model, it is possible to calculate pore water pressure generation by calculating irrecoverable volumetric strains during dynamic analysis. The void ratio in this model i s supposed to be constant, also it can be calculated as a function of volumetric strain and other parameters can be defined by void ratio (B.R. Khatibi et al, 2012). Martin et al. (1975) described initially the effect of cyclic loading on increase of pore water pressure as a result of irrecoverable volume contraction in the soil mass. In these situations, because the matrix of grains and voids is filled by water, the pressure of pore water increases (Itasca FLAC manual, 2008). It supply the following empiri cal equation that relates the increment of volume decrease vd to the cyclic shear shear strain (Itasca FLAC manual, 2008): vd = C 1 C 2 vd ) + (C 3 2 vd 4 vd ) Where C1, C2, C3 and C4 are constants. vd , in such a way that the increment in volume strain decreases as volume strain is accumulated.

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239 Pr vd is zero; this implies that the constants are related as follows: C 1 C 2 C 4 = C 3 . Martin et al. (1975) then go on to compute the change in pore pressure by assuming certain module and boundary conditions ((Itasca FLAC ma nual, 2008)). Later, Byrne (1991) presented a simpler equation which correspond irrecoverable volume change and engineering shear strain with two constants. In this model, a soil mass with liquefaction potential was modeled using (N 1 ) 60 parameter as a main factor to the Finn model, so all of the soil properties needed for the model were defined for the program by (N 1 ) 60 . ( ) = C 1 exp ( C 2 (( )) Where C 1 and C 2 are constants with different interpretations from above equation. In many cases, C 2 = 0.4C 1 , so the above equation involves only one independent constant; however, both C 1 and C 2 have been retained for generality (Itasca FLAC manual, 2008). As mention ed before, to the usual parameters (friction, module, etc.), the model needs the four constants for C1, C2, C3 and C4, or two constants for C 1 and C 2 . Martin et al. (1975) describes how four constants C1, C2, C3 and C4 may be determined from a drained cy clic test. Byrne (1991) notes that the constant, C 1 follows: C 1 = 7600(Dr ) 2.5 1 ) 60 Dr = 15(N 1 ) 60 1/2 Then: C 1 = 8.7 (N 1 ) 60 1.25 C 2 is then calculated from C 2 = 0.4C 1 in this case. Note that, as expected, the volumetric strain is larger for smaller values of the blow count (Byrne, 1991). Horita (1985) showed that the modelling simulated behavior of Monterey No.0/30 sand in a strain controlled undrained cyclic triaxial test and stress controlled undrianed cyclic triaxial test. A model for dilatant elasticity material during in undrained convent ional triaxial test with volumetric strain increment r =0 can be expressed as:

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240 = q = q = ( + + ) q = For triaxial conditions = = + ) Where and are axial and radial strain increments, respectively, are the principal effective stress increments and K, D 1 and D 2 are material constants, D 1 = 3G and D 2 = 6G, K is bulk modulus, G is shear modulus. If a pair of v v and shear strain component are selected for strain measures, the stress strain relationship becomes = q = q = + In the axisymmetric stress condition, in an undrained case, = 0, and 1 and D 2 Control shear induced pore water pressure which is not possible in an isotropic elastic material. F 1 = and F 2 = F 1 and F 2 are found to be constant in the undrained condition as F 1 = K ( ) F 2 = Horita (1985) showed that the slope of equi strain lines gradually increase and converges a critical slope at which the sand shows a large deformation and fails based on running

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241 compression and extension test on Monterey No. 0/30 sand. A parabolic relationship can be adopted to represent the above mentioned strain hardening behavior of Monterey No. 0/30 sand as M M o = Where M and M o are the slopes of an equi strain line and the zero strain line, respectively, is the axial strain and M f are materials constant. For compr ession test, M o = 0.75, M f = 1.43, and a = 0.00306, while in extension, M o = 0, M f = 1.0 and a = 0.0015. The equation for equi strain lines showed in simulation of drained conventional triaxial behavior. ) M where q is a deviator stress, p is strain line with the p axis, and M is the slope of the eui to be a constant equal to 3psi in both compression and extension. In undrained triaxial compre ssion, c a p = p o Where p o a is the axial strain, and b c , b e and U o are material constants. The simulation of undrained triaxial compression behavior requires three fundamental equations: (1) the equation for equi ) M (2) the equation for hardening rule for equi strain line: M = M o (3) the equation for shear induced pore water pressure in c a . When it simulated undrained triaxial compression behavior with given stresses, the equations are solved in terms of stresses, and the simulation process by fee ding stress increments is summarized as follows: 1) 2) q = q o 3) a a = where A = B=

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242 C = a 4) = b c a 5) 6) p = The simulation process of undrained triaxial extension tests by feeding stress increments is summarized as follows: 1) 2) q = q o 3) a a = where A = ) B= C = a b e q 4) = 5) 6) p = Horita (1985) run strain controlled undrained cyclic triaxial test on Monterey No. 0/30 sand to simulate the response of a sand to strain controlled undrained cyclic loading. The simulation consists of the following four steps: 1) Loading in compression 2) Unloa ding in compression 3) Loading in extension 4) Unloading in extension. Since in undrained condition specimen is 0.5. To account for nonlinear equi strain line s, a nonlinear equation: q = [1 (1+BP ) n ] where A,B, and n are determined from the shape of a equi strain line as shown in Figure 8.7. The sequential procedure for simulation of a strain controlled u ndrained cyclic loading behavior of saturated sand is presented in the flow chart in Figure 8.8. Horita (1985) showed that the response of Monterey No.0/30 sand in a stress controlled undrianed triaxial test, were conducted at a frequency of 0.5Hz and str ess ratio of 0.28, 0.32, and 0.37, can be simulated processes of stress controlled undrained cyclic triaxial behavior in both compression and extension and the undrained rebound behavior.

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243 As in the simulation of strain controlled undrained cyclic triaxial behaviors, the equations involved in the simulation of stress controlled cyclic triaxial behaviors are those for equi strian lines, hardening rules for equi strain lines, shear induced pore water pressure, and undrianed rebound behavior. The simulation pr ocedure is summarized in a flow chart in Figure 8.9. In the table 8. 1 , it showed that all para meter of trail 1 in Horital model have been used for simulating liquefaction behaviors of Monterey No. 0/30 sand. All parameter of Trail 2 showed in the table 8. 2. All parameter were measured from laboratory tests. In the Figure 8.10 1 and Figure 8.10 2 , it showed that the comparison cyclic triaxial test results wit h simulation by Horita model in excess pore pressure versus number of cycles to liquefaction. S olid curve s of figure 8.10 1 amd figure 8.10 2 are measured from cyclic triaxial test. D as h curve is simulat ed by using Horital model in Figure 8.10 1 and Figure 8.10 2 . In the trial 1, the simulated and measured curves had a gap, about 3 4 psi. However, i n the beginning of number of cycles of the trial 2 , the simulated and measured curves match ed well. T started to have lower value than the measure curve from cyclic triaxial test result after sixth cycle in the trial 2 . The different value is less 1 psi between measure ment and simulatio n curves in the t rial 2 . Soil sample was saturated Monterey No. 0/30 s and , and run it in the stress controlled undrained cyclic triaxal test. It was prepared relative density at 30% and run under stress ratio 0.4 and consolidation pressure 15psi, No. cycles to liquefaction =10 by Jungang Liu. The Fi between measure ment and simulat ion . Solid curve in Figure 8.11 showed the measurement from cyclic triaixl test, dash curve is simulated by Horita model. The measured and simulated curves as shown in Figure 8.11 match ed well.

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244 Figure 8.7 Curved Equi Strain Line (Horita, 1985)

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245 Figure 8.8 Flow Chart for Simulation on Strain Controlled Undrained Cyclic Traixial Tests (Horita, 1985)

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246 Figure 8.9 Flow Chart for Simulation on Stress Controlled Undrained Cyclic Triaxial Tests (Horita, 1985)

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247 Table 8.1 All parameters in Horita model fo r tr i al 1 (Masakuni, Horita, 1985) Figure 8.10 1 Comparison between Measured and Simulated Responses of Soil Samples in Generation of Pore Water Pressure. (Jungang Liu, 201 9 ) 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore pressure (psi) No, Cycles to liquefaction Measure Simulate Horita Model Parameters Description Compression Extension a Intersection of an equi strain q space 0.00087 0.0044 M f Ultimate slope of equi strain line 1.12 1 A Parameter for a curved failure line in n space 2.86 1.47 B Parameter for a curved failure line in n space 0.333 0.333 b Parameter for shear induced pore pressure 27 0.0256 F 1 Undrained rebound moduli 0.27 0.08 F 2 Undrained rebound moduli 0.0000659 0.0000185

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248 Table 8. 2 All parameters in Horita m odel for trial 2 (Masakuni, Horita, 1985) Figure 8.10 2 Comparison between Measured and Simulated Responses of Soil Samples in Generation of Pore Water Pressure. (Jungang Liu, 2019 ) 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore pressure (psi) No, Cycles to liquefaction Measure Simulate Horita Model Parameters Description Compression Extension a Intersection of an equi strain line q space 0.0006 7 0.003 4 M f Ultimate slope of equi strain line 1. 43 1 .2 A Parameter for a curved failure line in n space 2. 76 1.47 5 B Parameter for a curved failure line in n space 0.2 4 3 0.3 44 b Parameter for shear induced pore pressure 2 6 0.025 8 F 1 Undrained rebound moduli 0.31 0.0 78 F 2 Undrained rebound moduli 0.00006 62 0.00001 7 5

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249 Figure 8.11 Comparison between Measured and Simulated Responses of Saturated Monterey No. 0/30 Sand to Stress Controlled Undrained Cyclic Loading with stress ratio = 0.4, No. cycles to liquefaction =10 a) Effective Stress Path (Jungang Liu, 2019 ) -20 -15 -10 -5 0 5 10 15 20 0 2 4 6 8 10 12 14 16 q(psi) p'(psi) Measured in q vs p' -20 -15 -10 -5 0 5 10 15 20 0 2 4 6 8 10 12 14 16 q(psi) p'(psi) Simulated in q vs p'

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250 CHAPTER IX STATISTICAL MODELING OF LIQUEFACTION RESISTANCE Introduction Descriptive statistics and inferential statistics are the two major branches in the science of statistics and statistical methodology. Descriptive statistics deals with summary and description of data and inferential statistics concerns with analysis of sample data to make inferences about a large set of data a population, from which the sample is selected. Experimental research in engineering involves the use of experimental data a sample, to i nfer the nature of same conceptual population that characterizes a phenomenon of interest to the experimenter. One of the most important application of inferential statistics in engineering involves estimating the mean value if a response variable or predi cting some future values of the response variable based on knowledge of a set of related independent variables. A relationship used to relate a dependent (response) variable to a set of independent variables is generally referred to as, a regression model or a statistical model (Mendenhall and Sincich, 1991) Selection of Variables Initial Variable Selection In this research, uniform Monterey sand with six different percent of fines content, prepared at three various relative densities and consolidated isotr opically at the consolidation pressures of 15 psi and 30 psi. Cyclic triaxial test and cyclic hollow cylinder test were performed at cyclic stress ratios ranging from 0.2 to 0.4 and frequency of 0.5 Hz. Stress ratio in cyclic triaxial and cyclic hollow cyl inder tests causing initial liquefaction in 10 cycles, 30 cycles, 40 cycles and 50 cycles were chosen as dependent variables. The stress ratio causing initial liquefaction were determined from the test results presented in Chapter 4 and 5. Sixteen independ ent variables: 1 diameter corresponding to 50% finer in the particle size

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251 distribution, D 50 ; 2 . coefficient of uniformity, C u ; 3. coefficient of curvature C c ; 4. maximum void ratio e max ; 5. minimum void ratio e min ; 6. relative density, D r ; 7. stress ratio, S.R.; 8. Void ratio, e; 9. Number of cycles to liquefaction, N; 10. Fine content, F.C.; 11. Deviator stress, D.S. (in cyclic triaxial test), Shear stress S.S. (in cyclic hollow cylinder test); 12. Consolidation pressure, C.P.; 13. Plasticity index. PI; 14 . Overall void ratio, VRO; 15. Void ratio of fines, VRF; 16. Void ratio of sand skeleton, VRS were selected. Variable Reduction Introduction A total of one hundred and sixteen data were obtained in the variable reduction. Variable reduction is undoubtedl y one single most important task in the statistical model building. For the purpose of variable reduction, following procedures may be employed: 1. Examine the scatter diagrams and correlation matrix, 2. Study interaction and confounding terms in regressio n, 3. avoid collinearity, 4. Try different strategies for selecting variables during regression, 5. Perform factor analysis . Examine the Scatter Diagrams and Correlation Coefficient Matrix Examination of scatter diagrams of primary independent variables against dependent variables may reveal if high order terms of primary independent variables are needed. Also, necessity for the transformation of primary independent variables may be implicated. This will help in deciding some interaction terms to be cons idered initially. Correlation matrix of dependent variables and all independent variables including interaction terms and transformed terms can reveal strength of correlation between variables and help to decide what independent variables to be included in further analysis.

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252 Study Interaction and Confounding Term in Regression A number of options are available in using statistical testing to evaluate interaction for a given regression model. One approach is to test globally for the pressure of any kind of interaction and then, if significant interaction is found, to identify par ticular interaction terms of importance by using other test. A second way to assess interaction is to test for interaction in a hierarchical sequence, beginning with highest order term and proceeding sequentially to lower order terms if higher order terms are not significant. This should help to decide what interaction terms to be included in the regression. For primary independent variable that are not related to interaction terms, one can start with regression analysis using smaller number of independent variables and consider the rest of independent variables as potential confounders then exclude those which are not confounders to reduce number of variables. Avoid Collinearity When several interaction terms are included in a regression model, problem of collinearity may arise. The term collinearity is used to indicate that one of the independent variables is a linear combination of the others. Consequently, avoiding collinearity means reducing number of variables. To assess collinearity, the associated R 2 values based on fitting models for each of suspicious independent variables against the rest of independent variables are examined. If any of these multiple R 2 values equals 1.0, then a perfect collinearity exists among that particular set of variables. A R 2 >0.9 indicates define collinearity problem, while 0.8< R 2 <0.9 usually implies collinearity problem, a R 2 < 0.8 usually signals no collinearity problem. Many interaction terms may be excluded through collinearity elimination. The goal of variable reduc tion may be to eliminate collinearity, to simplify data analysis, or to obtain a parsimonious and conceptually meaningful summary of data. By achieving these goals, reduction of variable is

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253 in turn fulfilled. To determine the out set of independent variabl es to avoid collinearity, correlation matrix of all the independent variables can be established and following procedures may also be applied: 1. Delete the single independent variable, X j with the smallest tolerance (Tol j = 1 R j 2 ), thus defining the out set of size 1. R j 2 is the squared multiple correlation based on regressing X j against the rest of the independent variables. 2. Within the in set, delete the single independent variable with the smallest tolerance (this calculation ignores the variable deleted i n the previous step). 3. Compute the set of multiple squared correlations based on predicting each out set variable from the set of in set variables. 4. If the minimum multiple squared correlation from step 3 is too small, return the last deleted variable to t he in set and stop. Otherwise, deleting exactly one variable each time, repeat steps through 4 until the minimum multiple squared correlation is too small. Perform Factor Analysis Factor analysis is a multivariable method intended to explain relationship s among several difficult to interpret, correlated variables in terms of a few conceptually meaningful, relatively independent factors. Factors analysis may be conducted at the very beginning of the statistical study to help visualizing relationships among all the variables involved in the study. This certainly will help the process of variable reduction even if the results of the factor analysis are not used directly in the subsequent regression analysis. If the factors identified in the factor analysis ar e to be used in the regression, care must be taken regarding the meaning of each factor which is a weighted linear combination of the original variables. In engineering analysis, variables usually bear some obvious physical meaning. It would be highly desi rable if

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254 meaningful physical meaning can be attached to the factors resulted from factor analysis through the involving original variables. Otherwise confusion may arise concerning the physical meaning of the factors used in the regression and the regressi on results may not be convenient to use. Final Variable Selection In an effort to reduce variables for soils tested at three various relative densities of 30%, 45% and 60%. The correlation coefficient matrix of the dependent variable, SRCTT10, SRCTT30, SR CTT40, SRCTT50, SRCTT30 (Dr30) , SRCTT30 (Dr45) , SRCHCT10, SRCHCT30 and the sixteen chosen primary independent variables as shown in Table 9.1, 9.2, 9.3 9.4, 9.5, 9.6, 9.7, 9.8 were examined. Three independent variables, deviator stress in cyclic triaxial test, DS (cyclic shear stress in cyclic hollow cylinder test, CSS); fine content (decimal), FC; consolidation pressure, CP. were eventually selected for final statistical analysis.

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255 Table 9.1 Correlation coefficient matrix of the dependent variable, SRCTT10 and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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256 Table 9.2 Correlation coefficient matrix of the dependent variable, SRCTT30 and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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257 Table 9.3 Correlation coefficient matrix of the dependent variable, SRCTT40 and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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258 Table 9.4 Correlation coefficient matrix of the dependent variable, SRCTT50 and the sixteen chosen primar y independent variables (Jungang Liu, 2019 )

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259 Table 9.5 Correlation coefficient matrix of the dependent variable, SRCTT30 (Dr30) and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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260 Table 9.6 Correlation coefficient matrix of the dependent variable, SRCTT30 (Dr45) and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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261 Table 9.7 Correlation coefficient matrix of the dependent variable, SRCHCT10 and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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262 Table 9.8 Correlation coefficient matrix of the dependent variable, SRCHCT30 and the sixteen chosen primary independent variables (Jungang Liu, 2019 )

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263 Regression Model for Liquefaction Resistance Regression model for cyclic triaxial test A statistical model was formulated to predict the liquefaction resistance of soils containing different fine contents by using cyclic triaxial and cyclic hollow cylinder test results For cyclic tiaixal test results, stress ratio causing initial liquefacti on in 10 cycles, SRCTT10, was chosen as the dependent variable. Through the coefficient of correlation analysis, the independent variables were reduced from the initial sixteen to the final three: deviator stress, DS; fine content (decimal), FC; consolidat ion pressure, CP. Using the excel program, linear regression analysis was performed with one dependent variable SRCTT10 and three independent variables. The R 2 value for the regression equation is 1.0. A linear regression model involving DS, FC and CP as predictors was obtained as follows: SRCTT10=0.4+0.033*DS 7.03× *FC 0.027*CP Or SRCTT10=0.4+0.033*DS 0.027*CP (since coefficients of fine content was too small, one of independent variables was omitted.) For cyclic tiaixal test results, stress ratio causing initial liquefaction in 30 cycles, SRCTT30, was chosen as the dependent variable. Through the coefficient of correlation analysis, the independent variables were reduced from the initial sixteen to the final three: deviator stress, DS; fine content (decimal), FC; consolidation pressure, CP. Using the excel program, linear regression analysis was performed with one dependent variable SRCTT30 and three independent variables. The R 2 valve for the regression equation is 0.9. A linear regression model involving DS, FC and CP as predictors was obtained as follows: SRCTT30=0.305+0.02158*DS+0.033*FC 0.014*CP

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264 For stress ratio causing initial liquefaction in 40 cycles, SRCTT40 as the dependent variable, the same independent variables as it in 10 and 30 cyc les, the R 2 value for the regression equation is 0.965 . A linear regression model involving DS, FC and CP as predictors was obtained as follows: SRCTT40=0.2+0.033*DS+3.474× *FC 0.013*CP Or SRCTT40=0.2+0.033*DS 0.013*CP (since coefficients of fine content was too small, one of independent variables was neglected) Stress ratio causing initial liquefaction in 50 cycles, SRCTT50, and three independent variables. The R 2 value for the regression equation is 0.928. A linear regression model involving DS, FC and CP as predictors was obtained as follows: SRCTT50=0.3191+0.02*DS 0.0746*FC 0.0124*CP For soil sample prepared at relative density 30% and 45% in cyclic tiaixal test, stress ratio causing initial liquefaction in 30 cycles, SRCTT30 (Dr30) , SRCTT30 (Dr45) ,were selected as the dependent variable. Through the coefficient of correlation analysis, the independent variables were chosen: deviator stress, DS; fine content, FC; consolidation pressure, CP. Linear regression analysis was performed with one dependent variable SRCTT30 (Dr30) and three independent variables in data analysis of excel software. The R 2 value for the regression equation is 0.973. A linear regression model involving DS, FC and CP as predictors was obtained as follows: SRCTT30 (Dr30) = 0.223+0.017*DS+0.068*FC 0.008*CP Stress ratio causing initial liquefaction in 30 cycles, SRCTT30 (Dr45) , and three independent variables. The R 2 value for the regressi on equation is 0.942. A linear regression model involving DS, FC and CP as predictors was obtained as follows:

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265 SRCTT30 (Dr45) = 0.2883+0.0215*DS+0.1276*FC 0.0139*CP Regression model for cyclic hollow cylinder test For cyclic hollow cylinder test results, stress ratio causing initial liquefaction in 10 cycles, SR10, was decided as the dependent variable. The independent variables were picked: cyclic shear stress, CSS; fine content (decimal), FC; consolidation pressur e, CP after the coefficient of correlation analysis. Linear regression analysis was performed with one dependent variable SR10 and three independent variables in the excel program. The R 2 value for the regression equation is 0.95. A linear regression mod el involving C S S, FC and CP as predictors was obtained as follows: SR10=0.345+0.0595*C S S 0.001*FC 0.021*CP For cyclic hollow cylinder test results, stress ratio causing initial liquefaction in 30 cycles, SR30, was chosen as the dependent variable. Through the coefficient of correlation analysis, CSS, FC and CP were the independent variables. The R 2 value for the regression equation is 0.948. A linear regression model involving C S S, FC and CP as predictors was obtained as follows: SR30=0.299+0.0358*C S S 0. 0133*FC 0.011*CP

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266 CHAPTER X A REVISED PROCEDURE FOR EVALUATING LIQUEFACTION RESISTANCE OF SOIL WITH PLASTIC FINES Introduction In this study, the effect of fines contents on liquefaction resistance of a uniform Monterey No. 0/30 sand with different fines content was investigated using cyclic triaxial and cyclic hollow cylinder tests. A regression model to predict liquefaction resi stance of soils containing fines was formulated. Based on the procedures for evaluating field liquefaction potential of sand deposits as proposed by Seed and Idriss (1981) and Seed et al. (1985), the new findings of this study can be applied to better unde rstand the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils. Procedure by Seed, et al Seed and Idriss (1967) use the ratio of the earthquake induced cyclic stress ratios (CSR) with the cyclic resistance ratios (CRR) of the soil to evaluate the potential of soil liquefaction. situ parameter, such as CPT penetration resistance, SPT blow count, or shear wave velocity, V s . The procedure recommends the following equation for the evaluation of the earthquake induced cyclic stress ratio (CSR) with an equivalent uniform shear stress of 65% of the maximum shear stress that the soil experienced during the seismic wave propagation: CSR = 0. 65 where = maximum earthquake induced shear stress, = effective overburden stress at depth, z, where the subscript indicate the specific, earthquake magnitude (moment magnitude, M) and in situ effective overburden, . The choi ce of the reference stress level, 0.65 was selected by Seed and Idriss (1967) and has been in use since. The value of can be estimated from dynamic response analyses, but such analyses must include a sufficient number of input accelerati on time series and adequate site characterization details to be reasonably robust. Alternatively, the maximum shear stress can be estimated using the equation, developed as part of the Seed Idriss Simplified Liquefaction Procedure, which is expressed as, C SR = 0.65 r d

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267 v = vertical total stress at depth z, a max /g = maximum horizontal acceleration (as a fraction of gravitational acceleration) at the ground surface, and r d = shear stress reduction factor that accounts the reduction of shear stress from that of a rigid to flexible soil column in the soil deposit. C orrelation between cyclic shear stress ratio causing liquefaction in the field and normalized corrected (N 1 ) 60 standard penetration resistance of sand for the earthquake magnitude of 7.5 was presented as shown in Figure 10. 1. Fig. 10. 1 SPT case historie s of cohesionless soils with FC 35% and the NCEER Workshop (1997) curve and the recommended curves for both clean sand and for FC = 35% for M = 7½ and ' vo = 1 atm (I.M.Idriss and R.W.Boulanger, 2004) The stress based liquefaction analysis framework for soil includes four functions that describe fundamental aspects of dynamic site response, penetration resistance, and soil characteristics and behavior. These four functions, along with the major factors affecting each, are: 1) r d = f(depth; earthquake and ground motion characteristics; dynamic soil properties); 2) C N v ; DR; FC); 3) K v ; DR; FC) (shown in chap 3); 4) MSF = f(earthquake and ground motion characteristics; DR; FC) (shown in chap 3)

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268 Idriss (1999), in extending the work of Golesorkhi (1989), performed several hundred parametric site response analyses and concluded that, for the purpose of developing liquefaction evaluation procedures, the parameter r d could be expressed as, r d = ex 1.012 1.126 sin ( where z = depth below the ground surface in meters and the arguments inside the sin terms are in radians. Idriss and Boulanger (2010) summarize details r egarding the soil profiles and input motions used in developing these equations. The resulting variations of rd with depth and magnitude are shown in Figure 10. 2. Figure 10. 2 . Shear stress reduction factor, rd, relationship (I.M.Idriss and R.W.Boulanger, 2010) The C N relationship used was initially developed by Boulanger (2003b) based on: (1) a re v of 0.7 to 5.4 atm (Marcuson and Bieganousky 1977a, 1977b); and (2) result v of 0.2 to 20 atm using the cone penetration theory of Salgado et al. (1997a, 1997b) which was shown to produce v up to 7 atm.

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269 Idriss and Boulanger (2003, 2008) subsequently recommended that the D R dependence of the C N relationship could be expressed in terms of q c1Ncs or (N 1 ) 60cs as follows: C N = ( m = 1.338 0.249(q C1Ncs ) 0.264 m = 0.784 0.0768 with q c1Ncs limited to values betwe en 21 and 254 and (N 1 ) 60cs in these expressions. The values of C N calculated using this equation are presented in Figure 10. 3 (a) for a range of q c1Ncs and (N 1 ) 60cs values and for effective overburden stresses up to 10 atm, and are compared to the Liao and Whitman (1986) relationship in Figure 3 (b) for effective overburden stresses up to 2 atm. Figure 10. 3 . Overburden correction factor (CN) relationship for CPT and SPT penetration v ' /P a = 0 v ' /P a = 0 2 along with Liao and Whitman's (1986) relationship.

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270 Proposed Procedure Cyclic Resistance Ratio from SPT base d case history SPT based Case History Database from Idriss and Boulanger The individual SPT based liquefaction case histories data and key references are summarized by Idriss and Boulanger (2004, 2008) in Table 10. 1. The total number of case histories in the database is 230, of which 115 cases had surface evidence of liquefacti on, 112 cases had no surface evidence of liquefaction, and 3 cases were at the margin between liquefaction and no liquefaction. Idriss and Boulanger (2004) primarily used cases summarized in the databases compiled by Seed et al. (1984) and Cetin et al. (2 000, 2004), except that they excluded the Kobe proprietary cases that were listed in Cetin et al. (2004). The Fear and McRoberts (1995) database was also a helpful reference for many of the case histories. The updated database described in this report inc orporates the 44 Kobe proprietary cases which were provided by Professor Kohji Tokimatsu (2010), an additional 26 case histories summarized in Iai et al. (1989), and a small number of other additions. Data from the 1999 Kocaeli and Chi Chi earthquakes have not yet been incorporated. Idriss and Boulanger (2004, 2008) also primarily retained the values of critical depth, v v , and the product of the correction factors C E , C R , C B and C S listed by Seed et al. for the 1984 cases and by Cetin et al. for th e 2000 cases. Based on all 115 cases had surface evidence of liquefaction from Idriss and Boulanger (2004, 2008), it showed that the relationship between corrected SPT blow count (N 1 ) 60 and various fines content in Figure 10. 4. In the Figure 10. 4, 115 cas es proved that SPT blow count (N 1 ) 60 small than 30 had very high potential in soil liquefaction. Most of database is that soil had SPT blow count small than 25 with fines content with less than 35%. There are only 11 cases database that soil with fines con tent from 50% to 90% had liquefied in the Figure 10. 4. In the table 10. 2, it showed average of (N 1 ) 60 and average of cyclic stress ratio (CSR) at different percentages of fines content. There is a relationship between average of (N 1 ) 60 and different percen tages of fines content in cases histories database. The trend line of F igure 10. 5 showed the relationship between fines content with average of (N 1 ) 60 . The R 2 of trend line is 0.8673. In the figure 10. 5, it indicated clean fine content has the largest (N 1 ) 60 , (N 1 ) 60 decreases as fines contents increase. Beyond 15%, average of (N 1 ) 60 increas es with a further increase in fines content.

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271 In the F igure 10. 6, it indicated that the relationship between averages of earthquake induced cyclic stress rat io (CSR) and average (N 1 ) 60 with various fin es content (0%, 5%, 10%, 20%). It denoted that the10 percent of fines content has the lowest average of (N 1 ) 60 and smaller average of earthquake induce cyclic stress ratio. Figure 10. 4 . Case Histories from Idriss and Boulanger (all data for soil liquefy): (N 1 ) 60 versus FC (%) was created by Jungang Liu ( 2019 ) . 0 5 10 15 20 25 30 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 95 100 (N1)60 FC (%) FC vs (N1)60 from Idriss & Boulanger

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272 Figure 10. 5 . Case Histories from Idriss & Boulanger: Average (N1)60 versus FC (0% 25%) was created by Jungang Liu ( 2019 ). y = 0.002x 3 + 0.0947x 2 1.1976x + 15.148 R² = 0.8673 0 2 4 6 8 10 12 14 16 18 20 0 5 10 15 20 25 30 35 40 Average (N1)60 from Case Histories FC(%) Average (N1)60 from Idriss and Boulanger

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273 Figure 10. 6 . Case Histories from Idriss & Boulanger: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average (N1)60 with different FC (0% 20%) was created by Jungang Liu ( 2019 ). FC 0% FC 5% FC 10% FC20% 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0 2 4 6 8 10 12 14 16 18 20 Ave. Earthquake induced Cyclic Stress Ratio Ave. (N1)60 CSR vs Ave (N1)60 from Idriss&Boulanger

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274 Table 10.1 Summary of SPT based liquefaction case history data from Idriss and Boulanger (2004, 2008).

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275 Table 10.1 Summary of SPT based liquefaction case history data from Idriss and Boulanger (2004, 2008) (continue).

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276 Table 10.1 Summary of SPT based liquefaction case history data from Idriss and Boulanger (2004, 2008) (continue).

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277 Table 10. 2 Case histories data from Idriss and Boulanger: (N 1 ) 60 and CSR with different Fines Content data were created by Jungang Liu ( 2019 ). FC (%) 0 (0+) -5 (5+) -10 (10+) --20 (20+) -25 Total Samples 12 47 13 18 2 (N 1 ) 60 Ave. (N 1 ) 60 15.63 11.75 11.36 11.42 13.30 Standard Deviation 4.92 5.02 5.87 4.55 4.95 CSR Ave. CSR 0.30 0.26 0.25 0.28 0.29 Standard Deviation 0.15 0.10 0.09 0.12 0.03

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278 SPT based Cases History Database from Kohji Tokimatsu and Yoshiaki Yoshimi Many investigators have reported field evidence of soil liquefaction during strong earthquakes of which more than 70 case histories in Japan during 10 earthquakes as well as about 2 0 supplemental data outside Japan are available as shown in Table 10. 3. Tokimatsu and Yoshimi (1983) also primarily retained the values of critical water stable v v , SPT blow count N 1 (energy rod rati o of 78%), fine content (FC), median grain size ( D 50 ), coefficient of uniformity (Cu), clay content (CC), gravel content (GC) and calculated cyclic stress ratio (CSR) in Table 10. 3. The individual case histories and key references are summarized in Table 10. 3. The total number of case histories in the database is 95, of which 52 cases had surface evidence of liquefaction, 33 cases had no surface evidence of liquefaction, and 10 cases were at the margin between liquefaction and no liquefaction. In Figure 1 0. 7, it showed that SPT blow count N 1 versus different fines content from 0% to 65% from all 52 case database had liquefied. In Figure 10. 7, 52 case databases confirmed that it will be liquefied at SPT blow count N 1 small than 20. Additional, 80 percent of database is that soil had SPT blow count small than 20 with fines content with less than 35%. There are only 7 cases database that soil with fines content from 40% to 65% had liquefied in the Figure 10. 7. For four different percent of fines content (0%, 5 %, 10% and 20%), the average of N 1 , average (without maximum and minimum) of N 1 , average of calculating (CSR) and average (without maximum and minimum) of calculating CSR is shown in the Table 10. 4. In the figure 10. 8, one set of data is the average of N 1 with different percentage of fines content, another data is the average (without maximum and minimum) of N 1 with different percentage of fines content. Two sets of data had the similar curve. The curve of figure 8 had the same shape to the track of figure 10. 5 from Idriss and Boulanger (2004, 2008). The trend line of figure 10. 8 showed the relationship between fines content with average of N 1 . The R 2 of trend line is 0.9503.The clean fine content has the largest values N1 of 13.5 in the figure 10. 8. The N1 decreases with increase fines content until it reach 10 percent of fines content. After that, the trend is expected to go up with increase fines content. In the figure 10. 9, it showed that the SPT blow count always has the lowest value at the 10 percent of fines content, and also had the smallest earthquake induced cyclic stress ratio.

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279 From the figure 10. 5 and 10. 9, both data indicated that the clean fines content has the largest SPT blow count, and also 10% fine content of soil had the lowest SP T blow count. In the Figure 10. 10, two SPT based cases history databases from Idress, et. and Yoshimi, et. are in one plot. Although two databases had two different threshold fines content (one is 10%, another is 15%), both curves are the same trend sho wed in the figure 10. 10. Two databases showed that the clean fine content always had the largest values of SPT blow count. Case histories from different researchers showed that SPT blow count (N 1 ) 60 dropped with increasing percentage of fines content. Afte r threshold fines content, the (N 1 ) 60 started to increase with increasing fines content. Figure 10. 7 . Case Histories from Tokimatsu &Yoshimi (all data for soil liquefy): N1 versus FC (%) was created by Jungang Liu ( 2019 ). 0 5 10 15 20 25 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 95 N1 FC(%) N1 vs FC(%) From Tokimatsu & Yoshimi

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280 Figure 10. 8 . Case Histories from Tokimatsu &Yoshimi: Average (N1), average MM (N1) versus FC (0% 20%) was created by Jungang Liu ( 2019 ). y = 0.0264x 2 0.7721x + 13.141 R² = 0.9503 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Average N1 FC(%) FC(%) vs N1 case histories from Tokimatsu& Yoshimi Ave N1 Ave without MM Poly. (Ave without MM)

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281 Figure 10. 9 . Case Histories from Tokimatsu &Yoshimi: Average of Earthquake induced Cyclic Stress Ratio (CSR) versus Average N 1 with different FC (0% 20%) was created by Jungang Liu ( 2019 ). FC0% FC5% FC10% FC20% 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0.18 0.2 0 2 4 6 8 10 12 14 16 Earthquake induced cyclic stress ratio (CSR) Ave N1 CSR vs Ave N1 with different FC(%) from Tokimatsu & Yoshimi Ave CSR

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282 Figure 10. 10 . SPT blow count with different fines content (0% 25%) in case histories data from Idriss & Boulanger and Tokimatsu &Yoshimi (all data for soil liquefy) created by Jungang Liu ( 2019 ). 0 2 4 6 8 10 12 14 16 18 20 0 2 4 6 8 10 12 14 16 18 20 0 5 10 15 20 25 30 35 40 N1 (N1)60 FC(%) FC% vs N1, (N1)60 from SPT case histories Ave.(N1)60 from Idiss and Boulanger Ave N1 from Yoshimi Poly. (Ave.(N1)60 from Idiss and Boulanger) Poly. (Ave N1 from Yoshimi)

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283 Table 10.3 Summary of SPT based liquefaction case history data from Tokimatsu and Yoshimi (1983).

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284 Table 10.3 Summary of SPT based liquefaction case history data from Tokimatsu and Yoshimi (1983) (continue).

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285 Table 10. 4 Case histories data from Idriss and Boulanger: N 1 and CSR with different Fines Content data were created by Jungang Liu ( 2019 ). FC (%) 0 0 + -5 5 + -10 10 + -20 Total Samples 6 8 8 3 N 1 Ave. N 1 13.40 10.88 7.13 8.37 Ave without Max Min 12.88 10.65 7.53 Standard Deviation 4.03 2.69 3.10 2.20 Calculate CSR Ave. CSR 0.18 0.14 0.13 0.15 Ave without Max Min 0.18 0.14 0.13 Standard Deviation 0.05 0.05 0.02 0.03

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286 Cyclic Resistance Ratio from SPT based laboratory tests data Calculating SPT Blow Count (N 1 ) 60 Procedure Several early investigations were focused on the relationship between relative density and standard penetration resistance. Research regarding such relationship has been conducted (Gibbs and Holtz, 1957; Meyerhof, 1957; Skempton, 1986; Schultze & Menzenbac h, 1961; M.Cubrinovshi & K.Ishihara, 2001). In this study, it showed that the relationship between soil sample properties from laboratory tests such as relative density, confining stress, fines contents, and standard penetration resistance from field tests . There are six steps to calculate (N 1 ) 60 using soil sample properties from lab tests. Step 1: Calculate SPT blow count. a) For clean sand, Meyerhof (1957), the penetration resistance is assumed to increase with the square of the relative density and b e in direct proportion to the effective overburden pressure of the sand N = (17+24 ) where relative density expressed as a ratio, not a percentage. Skempton (1986) expressed in a general form N = (a + b ) Substituting = 98 kPa, the expres sion is reduced to = a + b where N 1 is the normalized penetration resistance to an overburden pressure of 98 kPa, i.e. 1 kgf/cm 2 . Note that in the original definition of Meyerhof the ratio or the parameter (a + b) was assumed to take a value o f 41. Schultze &Menzenbach (1961). Their data although presented as a logarithmic relationship, fit closely to the equation

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287 b) For soil containing fines For each undisturbed sample, the limiting void ratios (emax , emin) and relative density were evaluated by standard laboratory test procedures while the N1 value corresponding to the undisturbed sample was calculated from the known SPT blow count and sampling depth. The following empirical correlation by M.Cubrinovshi & K.Ish ihara (2001) between the SPT N value and Dr is derived N = v is given in kPa. It is important to note that, in this expression, the SPT blow count corresponds to an energy rod ratio of about 78 % of the theoretical free fall energy. Step 2 (only for soil containing fines): Relationship between e max , e min , and Fines Content Misko Cubrinovski and Kenji Ishihara (2002) showed the relationship between void ratio range and fines c ontent, (e max , e min ) is plotted against fines content in Figure 10. 11. A regression equation by Misko Cubrinovski and Kenji Ishihara (2002) relating maximum void ratio e max , minimum void ratio e min , and fines content, FC is following equation in the belo w. (e max , e min ) = 0.43+0.0086*FC, 0
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288 Step 3: Calculate N 60 N 60 can be converted by N values determined from the first step with a known or estimated ERr. Table 10. 5 showed that four countries have different rod energy ratios and hammers. N 60 = N Table 1 0. 5 Summary of Rod Energy Ratios (Skempton, A.W. 1986) Hammer Release ERr: % ERr/60 Japan Donut Tombi 78 1.3 Donut 2 turns of rope 65 1.1 China Pilcon type Trip 60 1 Donut Manual 55 0.9 USA Safety 2 turns of rope 55 0.9 Donut 2 turns of rope 45 0.75 UK Pilcon, Dando, old standard Trip 60 1 2 turns of rope 50 0.8 Step 4: Calculate (N 1 ) 60 In the field and laboratory tests described each of the sands (with one exception) is sufficiently uniform with regard to grain size and relative density to be treated as a unit, and the v =1 ton/ft 2 is found by direct interpolation. In ge neral, however, it is necessary to be able to estimate the N 1 value for any particular test, and this is done by means of the formula: N 1 = C N N The corresponding limits for C N are C N = Peck, Hanson&Thornburn (1974) C N = 0.77 log ( ) Liao and Whitman (1986) presented the currently held overburden correction, termed (N 1 ) 60 . The (N 1 ) 60 blow count is given as:

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289 (N 1 ) 60 = N 60 z is vertical effective stress where the sample was recovered. Calculation of SPT Blow Count based Laboratory Test Data In this study, one hundred fourteen cyclic triaxial and thirty seven cyclic hollow cylinder tests were performed to calculate SPT blow count (N 1 ) 60 by following the procedure in 10.3.2.1. After calculating (N 1 ) 60 using soil sample properties from laboratory tests, it showed that the relationship between standard penetration resistance and fines content. Calculation of SPT blow count based cyclic triaxial test d ata Ninety six cyclic triaxial tests were performed on the unifo rm Monterey No. 0/30 sand with six different percentages of fine content (5%, 10%, 15%, 25%, 35% and 45%) and plasticity index 20. Eighteen cyclic triaxal tests were run on the uniform Monterey No.0/30 clean sand. In Figure 10. 11, it indicated that the rel ationship between SPT blow count (N 1 ) 60 and different percentage of fines content in cyclic triaxial tests. In cyclic tiraxial test, soil samples prepared at three different relative densities of 30%, 45% and 60%, and also included six different percentage of fines content. Although three different relative densities, the SPT blow count (N 1 ) 60 had the biggest standard penetration resistance under clean sand. The (N 1 ) 60 decrease s with increase percentage of fines content and increase s with increas ing percentage of relative densities. In figure 10. 12, soil sample, prepared at 30% relative density in cyclic triaxal test, had the biggest SPT blow count (N 1 ) 60 when zero percentage of fine content. The SPT blow count (N 1 ) 60 decrease with i ncrease percentage of fines content until reach threshold fines contents at 25%. After 25% of fines content, SPT blow count (N 1 ) 60 started to go up at 35% of fines content.

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290 In the figure10. 13 and 10. 14, it showed that SPT blow count (N 1 ) 60 decreased with increase percentage of fines content under relative densities at 45% and 60% in cyclic triaxial test. In the figure10. 15, it showed that calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 30% under consolida tion pressure of 15psi with different fines content. Three different cyclic stress ratios of 0.2, 0.3 and 0.4 were applied in cyclic triaxial tests. The clean sand had the largest SPT blow count (N 1 ) 60 under three different cyclic stress ratios and consol idation pressure 15psi. The line of 25% of fines content located in the right side of curve in the figure 10. 9. However the line of fines content 35% found between the line of fine content 15% and the line of fine content 25%. It mean the lowest SPT blow co unt (N 1 ) 60 existed in fine content of 25% under three cyclic stress ratios. In the figure10.1 6, it showed that calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 30% under consolidation pressure of 30psi with different fines content. All lines of different fines content sited in similar to the figure 10. 9. Fine content of 25% had the lowest SPT blow count (N 1 ) 60 in cyclic triaxial test. In the figure 10. 17 and 10. 18, it showed that SPT blow count (N 1 ) 60 versus cy clic stress ratio on the relative density of 45% under consolidation pressure of 15 psi and 30psi with different fines content. In figure 10. 17 and10. 18, fine content of 5% always located in the right side of the curves, and gave the largest (N 1 ) 60 . Both figures showed that SPT blow count (N 1 ) 60 dropped with increase percentage of fines content. They indicated that SPT blow count (N 1 ) 60 increased with increase relative density. In the figure 10. 19 and 10. 20, it showed that SPT blow count (N 1 ) 60 versus cyclic stress ratio on the relative density of 60% under consolidation pressure of 15 psi and 30psi with

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291 different fines content. Both figures gave the same results from soil sample under relative densities of 30% and 45%.

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292 Figure 10. 11 Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35% , prepared at Dr= 30%, 45% and 60%. (Jungang Liu, 2019 ). Dr=30% Dr=60% Dr=45% 0 2.5 5 7.5 10 12.5 15 17.5 20 22.5 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on CTT under Dr= 30%, 45% & 60% E.S.=15psi &30psi From Jungang Liu

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293 Figure 10. 12 . Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 30% (Jungang Liu, 20 19 ) y = 0.0042x 2 0.2097x + 4.7839 R² = 0.9531 0 1 2 3 4 5 6 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=30% E.S.=15psi &30psi

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294 Figure 10. 13 . Cyclic Triaxial test results for calculating (N1)60 From FC=5% 35%, prepared at Dr= 45% (Jungang Liu, 2019 ). y = 0.0034x 2 0.2696x + 9.338 R² = 0.9765 0 1 2 3 4 5 6 7 8 9 10 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=45% E.S.=15psi &30psi

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295 Figure 10. 14. Cyclic Triaxial test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 60% (Jungang Liu, 2019 ). y = 0.0111x 2 0.7037x + 18.588 R² = 0.9669 0 5 10 15 20 25 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=60% E.S.=15psi &30psi

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296 Figure 10. 15 C alculating (N 1 ) 60 versus cyclic stress ratio from cyclic traixial test on relative density of 30% under consolidation pressure 15 psi with different fines content. (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=30% C.P.=15psi in CTT tests FC0ES15 FC5ES15 FC10ES15 FC15ES15 FC25ES15 FC35ES15

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297 Figure 10. 16 C alculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 30% under consolidation pressure of 30psi with different fines content (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=30% C.P.=30psi in CTT tests FC0ES30 FC5ES30 FC10ES30 FC15ES30 FC25ES30 FC35ES30

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298 Figure 10. 17 C alculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 45% under consolidation pressure of 15 psi with different fines content (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=45% C.P.=15psi in CTT tests FC5ES15 FC10ES15 FC15ES15 FC25ES15 FC35ES15

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299 Figure 10. 18 , calculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial t est on the relative density of 45% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=45% C.P.=30 psi in CTT tests FC5ES30 FC10ES30 FC15ES30 FC25ES30 FC35ES30

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300 Figure 10. 19 C alculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 60% under consoli dation pressure of 15 psi with different fines content (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=60% C.P.=15psi in CTT tests FC0ES15 FC5ES15 FC10ES15 FC15ES15 FC25ES15 FC35ES15

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301 Figure 10. 20 C alculating (N 1 ) 60 versus cyclic stress ratio in cyclic triaxial test on the relative density of 60% under consolidation pressure of 30 psi with different fines content (Jungang Liu, 2019 ). 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0 2 4 6 8 10 12 14 16 18 20 Cyclic Stress Ratio from Lab Calculate (N1)60 (N1)60 vs Stress Ratio on relative density=60% C.P.=30 psi in CTT tests FC0ES15 FC5ES30 FC10ES30 FC15ES30 FC25ES30 FC35ES30

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302 Calculation of SPT blow bount based cyclic hollow cylinder test d ata Twenty cyclic hollow cylinder tests were performed on the uniform Monterey No. 0/30 sand with five different percentages of fine content (5%, 10%, 15%, 25% and 35%) and plasticity index 20. Seventeen cyclic hollow cylinder tests were run on the unifor m Monterey No.0/30 clean sand. Calculating SPT blow count (N 1 ) 60 from cyclic hollow cylinder test results, it indicated that the relationship between (N 1 ) 60 and different percentage of fines content. In figure 10. 21, the trend is similar to the cyclic tria xial test under relative density at 30%. In figure 10. 22, the SPT blow count (N 1 ) 60 had the largest values about 20 at 0% of fine content. The value of SPT blow count (N 1 ) 60 start to drop after increasing fines content. Figure 10. 23 showed that two curves is the same trend to in cyclic triaxial test results under two different relative densities. The trend of figure 10. 24, it showed that the relationship between (N 1 ) 60 , cal culated from both cyclic triaxial and cyclic hollow cylinder test results, with different percentage of fines content. The R 2 of trend line is 0.9531. In the Figure 10. 24, soil samples prepared at Dr=30% with different percentage of fines content, run under consolidation pressure 15 psis and 30 psi on both laboratory tests. Although two database are from two different laboratory testing results, both curves are the same trend showed in the figure 10. 24. Both databases indicated that clean fines content had th e biggest value of calculated (N 1 ) 60. Both curves showed that the (N 1 ) 60 is dropping with increasing fines content, but after passing fines content of 25%, SPT blow count starts to increase. The R 2 of trend line in figure10. 25 is 0.9669. In the Figure 10. 25, soil samples prepared at Dr=60% with different percentage of fines content, run under consolidation pressure 15 psis and 30 psi in both cyclic triaxial and cyclic hollow cylinder test. Although two database are from two different laboratory testing res ults and different percentage of relative density, both curves are the same trend showed in the figure 10. 24 and 10. 25. Both databases indicated that clean fines content had the biggest value of calculated (N 1 ) 60. Both curves showed that the (N 1 ) 60 is droppi ng with increasing fines content, but after passing fines content of 25%, SPT blow count starts to increase.

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303 Figure10. 21 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 30% (Jungang Liu, 2019 ). y = 0.0044x 2 0.2166x + 4.8422 R² = 0.9606 0 1 2 3 4 5 6 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) Fc% vs Calculate (N1)60 on Dr=30% E.S.=15psi &30psi

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304 Figure 10. 22 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC = 0% 35%, prepared at Dr= 60% (Jungang Liu, 2019 ). y = 0.0107x 2 0.7204x + 19.3 R² = 0.974 0 5 10 15 20 25 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=60% E.S.=15psi &30psi in CHCT FC=0% 35%

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305 Fig ure 10. 23 Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35% , prepared at Dr= 30% and 60% (Jungang Liu, 2019 ). 0 5 10 15 20 25 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=30% & 60% E.S.=15psi &30psi in CHCT FC=0% 35% Dr=60% Cyclic Hollow Cylinder Test Dr=30% Cyclic Holloww Cylinder test Poly. (Dr=60% Cyclic Hollow Cylinder Test) Poly. (Dr=30% Cyclic Holloww Cylinder test)

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306 Figure 10. 24 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 30% (Jungang Liu, 2019 ). y = 0.0042x 2 0.2097x + 4.7839 R² = 0.9531 0 1 2 3 4 5 6 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 on Dr=30% E.S.=15psi &30psi from Jungang Liu's lab test results Cyclic Hollow Cylinder Test Cyclic Triaixal Test

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307 Figure 10. 25 Cyclic Triaxial test and Cyclic Hollow Cylinder test results for calculating (N1)60 From FC=0% 35%, prepared at Dr= 60% (Jungang Liu, 2019 ). y = 0.0111x 2 0.7037x + 18.588 R² = 0.9669 0 5 10 15 20 25 0 10 20 30 40 50 60 Calculate (N1)60 FC(%) FC% vs Calculate (N1)60 in CTT & CHCT on Dr=60% E.S.=15psi &30psi Cyclic Hollow Cylinder Test Cyclic Triaxial Test Poly. (Cyclic Triaxial Test )

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308 Comment on the New Procedure In this research, the field case histories data from Idriss .et al. (2004, 2008) and Youshimi. e t al. (1983) indicated that the soil with no fine content had the lar gest value of field test results SPT below counts in the figure10. 26. SPT below count, from field case histories, began to drop with increase percentage of fine content, but after passing threshold fines content, SPT below count tend to get bigger. Based entirely on laboratory test results, it also showed that SPT blow count decreases with increase fines content until threshold fines content. In the figure 10. 26, it showed that four different SPT be low counts, two from field case histories and another two from calculating SPT below count ba sed on laboratory test results, had the largest value when soil had no fines content. Threshold fines cont ent existed range from 10% to 25 % on SPT based case histo ries in the figure 10. 26. In laboratory test results in CTT and CHCT on Dr=30% & 60% soil samples, SPT blow count decreases with increase fines content until threshold fines content, ranges from 25% to 35% in figure 10.26. In the figure 10. 26, it showed that soils containing fines content from 10% 15% had the lowest number of SPT below count. For evaluating field liquefaction potential for soils containing fines content, the figure 10. 26 should be recommended. The results of the procedure should be applied to the evaluation of liquefaction resistance of soil containin g fines content from 5% to 25%. Specially, soil with different percentage of fines content should be paid more attention for eval uating liquefaction potentials.

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309 Figure10.26 . Calculating (N 1 ) 60 from cyclic triaxial and cyclic hollow cylinder tests on relative density 30% & 60% and consolidation pressure 15psi & 30psi, (N 1 ) 60 from case histories versus different fines content (Jungang Liu, 2019 ). 0 5 10 15 20 25 0 5 10 15 20 25 0 10 20 30 40 50 60 N1 from Yoshimi et al. (N1)60 from Idriss et al. and calculated from lab test results FC(%) Dr =60% in CTT and CHCT Dr =30% in CTT and CHCT Yoshimi & Tokimatsu Idriss & Boulanger

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310 CHAPTER XI SUMMARY, CONCLUSION AND RECOMMENDATIONS FOR FUTURE STUDIES Summary In this study, existing methods for liquefaction potential analysis and factors affecting liquefaction resistance of soils were reviewed. One hundred fourteen isotopically conso lidated undrain cyclic triaxial tests and thirty seven cyclic hollow cylinder test were performed to investigate the effect of fines content on liquefaction resistance of soils. A uniform Monterey No.0/30 sand and Leyden clay (was sieved through a #200 sie ve to remove any impurities) were involved. Fines with different consistency were prepared by mixing a sand and a silt at different percentage of composition. To compare and relate the soil liquefaction resistance found by cyclic triaixal and cyclic hollo w cylinder test results on uniform clean Monterey No.0/30 sand and soil sample with different percentage of fines content. The cyclic triaxial and cyclic hollow cylinder liquefaction resistance of soils were expressed in terms of liquefaction potential cu rves. From the liquefaction potential curves, stress ratio in cyclic triaxial and cyclic hollow cylinder tests causing initial liquefaction in 10 cycles, 30 cycles, 40 cycles and 50 cycles were chosen as dependent variables in the regression model. Three i ndependent variables of regression models, deviator stress in cyclic triaxial test, DS (cyclic shear stress in cyclic hollow cylinder test, CSS); fine content (decimal), FC; consolidation pressure, CP. were eventually selected for final statistical analysis. In addition to the evaluation of liquefaction potential, an excess pore pressure generation was performed from laboratory test results. Comparison of excess pore pressure generation between simulati is

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311 included in this research . The new findings of this study can be applied to better understand the field evaluation of liquefaction potential of soils containing fines and effect ively assess the liquefaction resistance of soils. Conclusions In this study, there are some following conclusions may be drawn: 1. The average of correction factor C r is 0.52 between cyclic triaxial and hollow cylinder test at the same relative densities o f 30 percent and 60 percent of uniform clean sand. 2. The average of correction factor C r is 0.557 between cyclic triaxial and hollow cylinder test at Dr =30% and 60% of samples with different fines. 3. At a constant overall void ratio and PI, the increase in f ines content up to 15% in the clean sand causes a decrease in liquefaction resistance. Beyond 15%, the further increases in fines content results in increase in liquefaction resistance. 15% is the threshold fines content observed from laboratory test resu lts in this study. 4. Threshold Fine Content is expected to occur when sand dominated behavior passes to fine dominated behavior. In this study, threshold fine content is approximately 15% based on cyclic triaxial and hollow cylinder test results. 5. Liquef action resistance of soils containing fines can be predicted by regression model which involves SRCTT10, SRCTT30, SRCTT40, SRCTT50, SRCTT30 (Dr30) , SRCTT30 (Dr45) , SRCHCT10, SRCHCT30 as indicators of liquefaction resistance while DS (CSS), FC and CP as indi cators of cyclic stress, fines content, and consolidation pressure. 6. Generation of excess pore water pressure due to cyclic stress is greatly affected by the fines content.

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312 7. e to cyclic stress. 8. After studying case histories databases and lab test results, the clean sand had the biggest number of SPT blow count, and also SPT blow count decreased with fines content increased. After passing threshold fines content, SPT blow count started to increase. For evaluating field liquefaction potential for soils containing fines content, soil with 10% 30% of fines content should be paid more attention. Recommendation for Future Studies In this study, there are some recommendations showed in below: 1. 2. Perform cyclic triaxial and hollow cylinder test on undistributed field soil samples. 3. Evaluating threshold fine content on different s oil types in laboratory tests. 4. For evaluating field liquefaction potential for soils containing fines content, Figure 10.26 need to be add more different soil types. It can help engineerings better understand the field evaluation of liquefaction potential of soils containing fines and effectively assess the liquefaction resistance of soils.

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313 REFERENCE S concrete buildings in Dagupan City due to liquefaction during the 1990 Philippine Netherlands, 147 152. Alarcon J. Geo tech. Engrg. , ASCE, 114(10), 1089 1109. siltgravel composites under Soil Dyn. Earthquake Eng., 18, 445 455. Geoenvir. Engrg., ASCE, 126(3), 208 217. tions under cyclic loading 126. wave 1015 1025. ASTM D 4253 , West Conshohocken, Pa. ASTM D 4254 , West Conshohocken, Pa. 99,West Conshohocken, Pa . 1025. evaluating NISTIR 6277, National Institute of Standards and Technology, Gaithersburg, Md. Bardet, J., Mace, N., and Tobita, T., (1999) "Liquefacti on induced ground deformation and failure." Report to PEER/PG&E.,University of Southern California. Denver. Beaty, M. H., and, P. M. Byrne (2011), UBCSAND constitutive model. Version 904aR,

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319 Kokusho, T.: Liquefaction strengths of poorly graded and well graded granular soils investigated by lab tests, 4th International Conference on Earthquake Geotechnical Engineering, Thessaloniki, 159 184, Springer, 2007. Ku, Chih Sheng, Der Her Lee, and Jian Hong Wu. "Evaluation of Soil Liquefaction in the Chi Chi, Taiwan Earthquake Using CPT." Soil Dynamics and Earthquake Engineering 24.9 10 (2004): 659 73. Elsevier. Web. eoenvironmental Engineering, June, 557 567. GED, ASCE, Vol. 100, No. GT4, April, PP. 387 406. Liu Hsing Cheng Liu Hsing Cheng and Nien of a uniform Fin Earthquake Engineering, 251 260. Liu Hsing Cheng and Nien SITU Liquefaction Resistance of a uniform Fine Sand with Plastic National Conference on Earthquake Engineering, 261 270. Liu Jungang and Nien Conference on Transportation Infrasturcture and Materials, QINGDAO, China. Makia, A. (2013), Evaluation of the UBC3D PLM constitutive model for prediction of earthquake induced liquefaction on embankment dams, MSc Thesis, TU Delft. Madabhushi, S.P.G., Knappett, J.A. and Haigh, S.K., (2009). Design of pile foundations in liquefiable soils, Imperial College Press, ISBN 1848163622. Martin, G.R., Liam Finn, W.D., and Seed, H.B. (1975), "Fundamentals of liquefaction under cyclic loading", Journal of Geotechnical Engineering Div. , ASCE, 101, GT5, pp423 437. Master Thesis, University of Colorado Denver. Mehmet Murat Monkul Influence of silt size and content on 942. Meyerhof, G.G. ( 1957 ) . Discussion on Research on determining the density of sands by

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320 penetration testing. Proc. 4th Int. Conf. on Soil Mech. and Found. Engrg ., Vol. 1: 110. Misko Cubrinovski, Jonathan D. Bray, Merrick Taylor, Simona Giorgini, Brendon Bradley, Soil Liquefaction Effects in the Central Research Letters, Volume 82, Number 6 2011. Misko Cubrinovski and Kenji Ishihara (2002). Maximum and Minimum Void Ratio Characteristics of Sands Soi l and Foundations Vol. 42, No. 6, 65 78, December 2002. Behavior of Shallow Foundations on Liquefiable Sand, Using Video Processing onic Journal of Geotechnical Engineering Vol.16, 945 960. based probabilistic assessment of seismic soil liquefaction California at Berkeley, Berkeley, CA. EE 001, National Academy Press, Washington, D.C. www.novotechsoftware.com Park, I.J., and Desai, C.S. (2000). Cyclic behavior and liquefaction of sand using disturbed state concep t, J. Geotech. Geoenv. Eng., ASCE, 126, 9. Petalas, A., V. Galavi, and R.B.J. Brinkgreve, (2012), "Validation and verification of a practical constitutive model for predicting liquefaction in sands". Proceedings of the 22"'' European young geotechnical engineers conference, Gothenburg, Sweden, pp. 167 172. Pradhan, S.K., and Desai, C.S. (2006). DSC model for soil and interface including liquefaction and prediction of centrifuge test, J. of Geotech and Geoenviron. Eng., ASCE, 132, 2 , 214 222. Polito, C. P., and Martin, J. R. (2001). Effects of nonplastic fines on the liquefaction resistance 415. Puebla, H., M. Byrne, and, P. PhilHps, (1997),"Analysis of canlexliquefaction embankments prototype and centrifuge model s", Canadian Geotechnical Journal 34, pp. 641 657. DUT, the Netherlands, 2013

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322 Fines Content of Mixtures of Sand and Non Vol.29, No.2 Skempton, A.W. ( 1986 ) . Standard penetration test procedures and the effects in sands of overburden pressure, relative density, particle size, ageing and overconsolidation. Geotechnique 36(3):425 447. . 3rd International Conference on Recent Advances in Geotechnical Earthquake Engineering and Soil Dynamics, St. Louis, Vol.1 249 252, 1995. 324 . S.S.C. Liao and R.V. Whitman, 1986, Overburden Correction Factors for SPT in Sand: Journal of Geotechnical Engineering , A.S.C.E., v. 112:3, p. 373 377. Non/Low Journal of Geotechnical and Geoenvironmental Engineering, 138(6), 747 756. Tatsuoka, F., Iwasaki, T., Tokida, K., Yasuda, S., Hirose, M., Imai, T. & Konno, M. (1978). A method for estimating undrained cyclic strength of sandy soils using standard penetrati on resistance. Soils and Foundations 18(3): 43 58. J. Geotech. Geoenviron. Eng. , Vol. 124, No. 6, pp. 479 491. Thevanayagam liquefaction pore Advances in Geotechnical and Earthquake Engineering, Paper No. 4.28. Screening and Soil Dyn. Earthquake Eng. , Vol. 22, No. 9 12, pp. 1035 1042. Tokimatsu , K., and Yoshimi, Y. (1983). "Empirical correlation of soil liquefaction based on SPT N value and fines content." Soils and Foundations, 23(4), 56 74. Journal of Geotechnical and Geoenvironmental Engineering , 129(4), 315 322. and SPT based probabilistic Japan Workshop on Earthquake Resistant Des. of Lifeline Facilities and Countermeasures Against Liquefaction Technical Rep. MCEER 99 0019, Multidisciplinary Center for Earthquake Engineering Research, Buffalo, N.Y., 69 86.

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323 Tsegaye, A., (2010), "Plaxis liquefaction model", external report. PLAXIS knowledge base: www.plaxis.nl essure Changes in Sands under reconstituting method Geotech.Testing J. , 22(3), 187 195. Geotechnical Engineering Journal of Ihe SEAGS &AGSSEA Vol. 44 No.3 Department of Civil & Environmental Eng. Massachusetts Institute of Technology, February 10 th , 2012. Vucetic, M., and Dobry, R. (1986). Pore pressur e build up and liquefaction at level sandy sites during earthquakes. Research Rep. CE 86 3, Dept. of Civil Engineering, Rensselaer Polytechnic Institute, Troy, NY. Wilson, D.W., Boulanger, R.W., and Kutter, B.L. (1997). Soil pile superstruc ture interaction at soft or liquefiable soil sites Centrifuge data report for CSP02. Rep. No. UCD/CGMDR 97/05, Ctr. for Geotech. Modeling, CEE Dept., Univ. of California, Davis. Wu, J., Seed, R.B., and Pestana, J.M. (2003). Liquefaction tr iggering and post liquefaction deformations of Monterey 0 30 sand under unidirectional cyclic simple shear loading. Geotechnical Engineering Research Rep. No. UCB/GE 2003/01, Univ. of California, Berkeley, CA. Xia Gao, Ling X, Soil Pile Bridge Structure Interaction in Liquefying 1009 1017 . Can. Geotech. J. , Ottawa, 34, 905 917. 324. Yamamuro, Soil Dyn. Earthquake Eng. , Vol. 24, pp. 751 760. Study of Effects of Fines Content on Liquefaction GeoShanghai 2010 International Conference Soil Dynamics and Earthquake Engineering, 272 277.

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324 ted sand during 38. ken Nanbu Association for Earthquake Engineering, Vol.4, No.3, 2004. Found., 41(4), 25 45. Liquefaction and Deformation of Silty and Fine Grained Soils 1013 1020 Youd 1996 NCEER and 1998 NCEER/NSF Workshops on evaluation of liquefaction resistance of 313. Y.P V Confining stress and static shear effects in cyclic Can. Geotech. J. 2001( 38) : 580 591. induced ground s Canadian Geotech. Journal , 39(5), 1168 1180. Liquefaction Potential of Silt Conference on Earthquake Engineering, Tokyo Kyoto, Japan, August, Vl. III. 237 242.

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325 APPENDIX A A. Field Methods for Soil Liquefaction Resistance Evaluation A.1. Other Techniques A.1.1 Becker Penetration Test and Large Penetration Test Becker penetration test (BPT) and large penetration test (LPT) have been used in soils with large particles (e.g., gravels and cobbles) that can interfere with the accuracy of SPTs and CPTs or even preclude their use. The BPT uses a double acting diesel pile hammer to drive into the ground a 168 mm diameter, 3 mlong double walled casing with a closed b it. The BPT test provides a continuous driving record, from which the blow count is the number of hammer blows required to drive the casing each 300 mm (1 ft) into the ground. The LPT is similar to an SPT, except that it uses a larger split spoon sampler a nd a larger hammer to drive it. The BPT depends on a number of factors that affect the energy delivered to the casing Correlations between BPT and SPT values in sand deposits are used to convert BPT blow counts into equivalent SPT N values for use in liquefaction analyses. Attempts have been made to further standardize the BPT and better understand its mechanics, but significant concerns remain about its repeatabi lity and general interpretation. More details about the BPT procedures and related issues are provided by Harder (1997) and Sy (1997). Several different LPTs have been developed around the world that have sampler outer diameters of 7.3 14 cm (as compared w ith 5.1 cm for the SPT) and hammer potential energies of 1.2 5.9 times the potential energy for the SPT hammer. Penetration resistances from LPTs have been correlated with those from SPTs, so the LPT values can be converted into equivalent SPT N 60 values f or use in liquefaction evaluations. Daniel et al. (2003) showed that wave equation

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326 analyses of the different penetration tests provided a rational means for assimilating various empirical LPT SPT correlations. They further noted the importance of energy me asurements for obtaining reliable LPT penetration resistances. A.1.2 Seismic Wave Velocity Chen Yunmin and Chen Pen Peng (2005) determined the response of reconstituted laboratory soil specimens to a given low amplitude P wave excitation, demonstrating a r elationship between acoustic signature so measured and liquefaction resistance. The study was aimed at evaluation of liquefaction potential in marine deposits where sampling is particularly difficult and a data base exists for acoustic response. Seismic w ave velocities (P wave and shear, S wave) are routinely determined through field geophysical surveys to obtain input for dynamic response analysis (Department of the Army, 1999). Measured shear wave velocities can be normalized to a standard effective ove rburden pressure of 1 ton/ft 2 (96 kPa) by V sl = V s vo ) 1/n (2.1) vo is in tons/ft 2 and n is taken as 3 (Tokimatsu et al., 1991) or 4 (Finn, 1991; Kayen et al., 1992). Stokoe et al. (1988) used the cyclic strain approach and equivalent linear ground response analyses to explore the relationship between peak ground surface acceleration (for stiff soil site conditions) and shear wave velocity. The results were used to develop bounds for the conditions under which liquefaction could be expected; the results agreed well with observed behavior in two earthquakes in the Imperial Valley of California (Figure 2.27). Tokimatsu et al. (1991) used the results of laborat ory tests to develop curves showing the CSR

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327 required producing cyclic strain amplitude of 2.5% in various numbers of cycles as a function of corrected shear wave velocity (Figure 2.28). The observation that the shear wave velocity of sand is insensitive to factors (e.g., soil fabric, over consolidation ratio, prior cyclic straining) that are known to influence liquefaction resistance suggests that shear wave velocity measurements alone may not be sufficient to evaluate the liquefaction potential of all soil deposits (Jamiolkowsky and LoPresti, 1992; Verdugo, 1992). Tokimatsu, Yoshimi and Uchida (1996) proposed a method to evaluate in situ liquefaction resistance of dense sands that may eventually prove adaptable to other soils, wherein: (1) shear wave veloci ties are determined by geophysical survey; (2) high quality samples are obtained by in situ freezing; (3) laboratory initial shear modulus, G max , is determined by low amplitude cyclic shear testing (type of equipment unspecified) and compared to that calcu lated from field shear wave velocity; (4) laboratory G max is adjusted (increased) by application of low amplitude (equipment again unspecified) preshearing until field and laboratory values match, and (5) cyclic triaxial tests are performed to measure liqu efaction resistance of thawed specimens. Adjusted specimen liquefaction resistance is claimed to represent in situ behavior, Stokoe, et al. (1988) developed charts relating shear wave velocity to maximum surface acceleration, a max , that predict liquefaction potential in clean sands (e.g., Figure 2.29).

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328 Figure 2.27 Chart for evaluation of liquefaction potential from shear wave velocity and peak ground acceleration ( 0 cycles). (After Stokoe eta., 1988.) Figure 2.28 Correlations between cyclic stress ratio required to produce cyclic strain amplitude of 2.5% in clean sand and shear wave velocity. (After Tokimatsu eta., 1991.)

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3 29 The shear velocity of the soil can also be used to determine the factor of safety against liquefaction. The shear wave velocity can be measured in situ by using several different geophysical techniques, such as the up hole, down hole, or cross hole methods. Other methods that can be used to determine the in situ sh ear wave velocity include the seismic cone penetrometer and suspension logger (Woods 1994) The shear wave velocity is corrected for the overburden pressure by using the following equation (Sykora 1987, Robertson et al. 1992): V sl = V s C v = V s vo ) 0.25 (2.8) Where V sl = corrected shear wave velocity C v = correction factor to account for overburden pressure. C v vo ) 0.25 vo = vertical effective stress kPa V s = she ar wave velocity measured in field

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330 Figure 2.30 Relationship between cyclic resistance ratio and corrected shear wave velocity for clean sand, silty sand, and sandy for M= 7.5 earthquake (From Andrus and Stokoe (20 00), required with permission of the American Society of Civil Engineers.)

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331 Fig2.31. Curves with new CSR V s 1 database (modified from Kayen et al. 2004) Figure 2.30 showed that the corrected shear wave velocity V s1 , and then by intersecting the appropriate fines content curve, the cyclic resistance ratio is obtained. The first is the global Vs database presented by Kayen et al. (2004). Fig.2.31 presents 60% of the global V s1 data and the corresponding probability curves, and the present lower bound CRR V s1 curve and that of Andrus and Stokoe (2000) are also plotted. The present curve separates all the liquefied case data properly in a slightly conservative way, and approximat ely corresponds to that of PL=0.05 in the range of 100 m/ s200 m/ s.

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332 An advantage of using the shear wave velocity to determine the factor of safety against liquefaction is t hat it can be used for very large sites where an initial evaluation of the liquefaction potential is required. Disadvantages of this method are that soil samples are often not obtained as part of the testing procedure, thin strata of potentially liquefiabl e soil may not be identified, and the method is based on small strains of the soil, whereas the liquefaction process actually involves high strains. In addition, as indicated in Fig. 2.30, there are little data to accurately define the curves above a CRR o f about 0.3. Furthermore, the curves are very steep above a shear wave velocity of 200 m/s, and a small error in measuring the shear wave velocity could result in a significant error in the factor of safety.

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333 Figure 2.29 Chart to predict liquefaction i n clean sands from shear wave velocity and maximum acceleration (Stokoe, et al. 1988)

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334 A.1.3. Other Techniques A number of additional in situ testing techniques show promise as tools to assist in site characterization for liquefaction potential evaluation. Any or all in situ techniques may someday prove useful in the assessment of cyclic strength of fine grained so ils, since the soils of concern are difficult to sample. The self boring pressure meter was used to evaluate liquefaction potential of sand through correlation with the dilation angle parameter. Dilation angle, defined as the inverse sine of the slope of a volume expansion versus shear strain curve, may be measured either from drained laboratory triaxial of simple shear tests or from in situ pressure meter tests. Pilot tests on a hydraulic fill dam yielded reasonably similar estimation of liquefaction resi stance from SPT blow count based and pressure meter based techniques. Electrical resistivity and conductivity geophysical survey methods have been applied to characterize in situ properties using either surface or borehole sensor arrays (Department of the Army, 1999). They studied electrical anisotropy of soil deposits, developing a structural index that may correlate to cyclic strength. Erchul and Gularte (1982) investigated densification in liquefying sand deposits in the laboratory using electrica l resistivity; they proposed extending the method to evaluate field deposits and monitor compaction efficiency.

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335 APPENDIX B B. Factors Affecting Liquefaction Resistance of Soils B.1 Effects of Laboratory Factor B.1.1 Specimen Preparation Method The effect of sample preparation on the cyclic strength of soils was presented by Ladd (1974). An electro hydraulic closed loop loading system was used in his tests. Samples were prepared by two different specimen preparation methods to investigate their e ffect on the cyclic strength of three materials with different gradation. The specimen prepared by the wet tamping method was found to be always stronger than the specimen prepared by the dry vibration method. Silver, et al. (2000) also proved that the cyclic strength of the specimen prepared by using the dry vibration method was on the order of half the strength of the specimen prepared by using the wet tamping method. The cyclic strength of the specimen prepared with the dry method did not increase si gnificantly with increasing stress ratios. Mulilis, et al. (1977) presented the most comprehensive studies regarding specimen preparation effects on the cyclic triaxial test. Six procedures with different specimen preparation methods were used in the stres s controlled cyclic triaxial tests. The effect of the method of sample preparation on the liquefaction characteristics was found to be significantly different. Differences in the cyclic stress ratio causing initial liquefaction of Monterey No. 0 /30 Sand w ere found to be in the order of 100%. Generally speaking, the weakest specimens were formed by pluviating sand through air, while the strongest specimens were those formed by vibrating the soil in a moist condition.

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336 Silver, et al. (2000) compared the cycli c strengths of specimens prepared by moist vibration, moist tamping, dry tamping, and dry vibration. The same conclustion as that of Mulilis, et al. (1977) was found. Furthermore, Mulilis (1978) presented the data obtained on specimens of Monterey No. 0/30 Sand prepared by the moist rodding and the dry rodding methods. An increase of cyclic strength of approximately 50 percent at 10 cycles to cause initial liquefaction was noted. In the same publication, the effect of tamping foot size was also examined. Ho wever, no significant effect on the tamping foot size was found. B.1.2 Reconstitution versus Intact Specimens As specimen preparation procedure had a strong influence on cyclic triaxial strength (Mulilis, 1978), dilemma may arise as to what reconstitutio n method should be adopted for comparison here. Limited data using moist tamping and pluviation device through water to reconstitute specimens have shown that undisturbed specimens were slightly stronger than reconstituted specimens (Ishihara et al., 1978; Mulilis et al., 1978). It should be noted that cyclic triaxial strength of undisturbed specimens are subjected to such factors as degree of in situ cementation and amount of disturbance during sampling. B.1.3 Load Wave Forms It has been found that wave forms of cyclic loading affect liquefaction resistance. Mulilis et al. (1978) compared the effects of rectangular, triangular, and sine wave loading as shown in Figure C.1. In Figure C.1, the order of increasing strength was rectangular, triangular, and si ne, with triangular and sine wave loading strengths being 13 and 30% stronger than rectangular wave loading, respectively. Results of similar trend were also reported by other researchers (Lee

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337 and Fitton, 1989; Silver et al., 2000). The effect of loading w ave form has been extensively studied by researchers and the results from these studies are quite similar. Silver, et al. (2000) performed a series of cyclic triaxial tests using three different wave forms; (1) a sine wave; (2) a square wave with a very r apid rise time; and (3) a square wave with a degraded rise time whereby the unloading and loading portions of the wave did not have an instantaneous change in velocity. Results from these tests show that the cyclic soil strength is significantly affected b y the shape of the loading wave. Specimens tested using a fast rise time square wave showed cyclic strength values approximately 15% less than those tested using a sine wave loading or a degraded square wave pattern. Examination of the pore pressure respon se recorded during a sharp square wave loading indicated that the instantaneous changed in velocity caused a stress wave to propagate through the specimen. This stress wave was reflected in the form of pore pressure spikes. The more rapid buildup of pore p ressure associated with the sharp square wave caused the sample to liquefy in a fewer number of cycles. It was observed that if the rise time in the rectangular wave form was degraded such that the wave form did not have an instantaneous change of veloci ty in either the loading or unloading portion of the cycle, then the strength of specimens which were tested using thedegraded wave form was approximately the same as that of specimens which were tested using the sine wave form. Due to the rapid jump in po re pressure associated with severe square wave loading, Silver recommended that a degraded square wave with a rise time of approximately 10% of the loading period or a sine wave loading be used in cyclic triaxial testing.

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338 Number of cycles Figure C.1 Effect of loading wave form on cycles to initial liquefaction for moist tamped Specimen B.1.4 Frequency on Cyclic Strength The effect of frequency over a range of 1/12 to 60 Hz on cyclic strength has been inconclusive with some researchers (Lee and Fitton, 1989; Lee and Focht, 1995) reported that slower loading frequencies produced slightly (< 10%) lower strength while others (Wulilis, 1975; Wong et al., 1975) reported otherwise. A study on the effect of frequency ranging from 0.00011 to 1 Hz showed that below 0.01Hz. Cyclic strength was independent of frequency effect while above 0.01 Hz, cyclic strength tend to increase with increasing frequency (Samuelson, 1981).

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339 B.1.5 Specimen Size A previous study concluded a height to diameter ratio of 2 is usually required. Lee and Fitton (1989) reported little effect on cyclic strength between specimen size of 1.4 and 2.8 inches in diameter. Larder (1999) however, reported lower liquefactio n resistance in specimens with 2.8 inches diameter than those with 1.4 inches diameter due to the effect of membrane penetration. Another study by Wang et al. (2002) involved specimen size of 2.8 and 12 inches in diameter showed similar membrane penetratio n effect. B .1.6 Frictionless Caps and Bases The cap and base friction of the triaxial specimen might be different for sample of different diameters. The effect of caps and bases friction on cyclic strength has been reported to be insignificant (Mulilis, 1 975). B .1.7 Membrane Compliance To minimize this effect, a relatively thick membrane was used to reduce the amount of initial penetration into the irregular sample surfaces. Martin, et al. (1978) investigated the effect of system compliance on uniform sands. They concluded that membrane compliance affected well graded samples. In addition, samples containing a small proportion of gravel samples. In addition, samples containing a small proportion of gravel would produce a relatively large void on the sample surface, leading to a large increase in the apparent resistance to liquefaction. When pressure during consolidation is applied to a sample through a rubber membrane, the mem brane deforms and is pushed into the pore spaces between the grains. This results in

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340 expulsion of some pore water from the sample, without a change in void ratio of the sample. Thus the measured volume change during consolidation must be corrected for memb rane penetration when void ratio is calculated. There are a number of theoretical studies (Molenkamp and Luger,1981; Baldi and Nova, 1984; Kramer et al., 1990) summarized by Ali et al., (1995) suggesting the form of the equation for membrane penetration. F or practical purposes, membrane penetration can be quantified in terms of a normalized membrane penetration: m m / A s log (p 1 /p 2 ) m = normalized membrane penetration m = volume change due to membrane penetration As = sample area covered by the membrane (2 rh for a cylindrical sample) p 1 p 2 = net pressure acting across acting the membrane before and after the volume change m is primarily dependent on grain size, assuming other factors such as membrane thickness and modulus are c ontent. B.1.8 Relative Density In one of the earliest laboratory cyclic triaxial study, Seed and Lee (1966) concluded that void ratio of a saturated sand strongly affected its liquefaction resistance the higher the void ratio or the lower the relative density, the more easily liquefaction will occur. Lee and Seed (1967) reported that cyclic stress required causing initial liquefaction increased linearly to approximately 60% relative density. Other study showed that the stress ratio to cause liquefactio n in 10 cycles is linear with relative density to approximately D r = 70% (Mulilis, 1975). The paramount importance of relative density as a parameter of liquefaction resistance

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341 was evidenced in various empirical correlations based on observations during pr evious earthquakes for the evaluation of liquefaction potential (Kishida, 1969; Castro, 1975; Seed and Idriss, 1981; Tokimatsu and Yoshimi, 1983). In these correlations, SPT N value, which has been shown to relate to relativity density (Gibbs and Holtz, 19 57) of soil, is invariably used as an indicator of soil strength liquefaction. In all these correlations, the lower the SPT N values or the lower the relative density, the lower the liquefaction resistance. B.1.9 3 ) Seed and Lee (1966) reported that liquefaction resistance of a saturated sand was affected by the confining pressure acting on the sand the lower the confining pressure the more easily liquefaction will develop. The effect of confining pressure on liquefa ction resistance of soils as concluded above is consistent with the fact that soil strength increases with confining pressure. However, confusion may arise if cyclic stress ratio instead of absolute cyclic stress amplitude is used to designate intensity of cyclic loading. In using equivalent uniform stress cycle concept (Seed et al., 1975; Annaki and Lee, 1977) for soil liquefaction analysis, it is convenient to express in situ cyclic loading in terms of cyclic stress ratio which is a ratio of cyclic shear stress amplitude to effective overburden pressure. In a one dimensional simplification, a magnitude of earthquake induced cyclic shear stress in a soil is in direct proportion to effective overburden pressure it is subjected to (Seed and Idriss, 1967). In laboratory triaxial condition, effective overburden pressure in the field can be simulated by effective confining pressure if in situ coefficient of lateral earthquake, K is equal to unity. Applicability of laboratory triaxial condition for different in si tu K values was discussed by Seed and Peacock (1970). Therefore, in the event of an earthquake shaking, soils under higher effective overburden pressure or effective confining pressure will in general experience higher shear stress amplitude and vice versa . Due to this

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342 confining pressure dependency, stress ratio, being a confining pressure normalized parameter is apparently a better indicator for liquefaction resistance under earthquake loading. Observations in laboratory have confirmed that cyclic stress r atio required to cause liquefaction decreases with increasing confining pressure (Castro and Poulos, 1976; Mulilis et al., 1977). It can be concluded that when cyclic stress ratio is used to designate cyclic loading intensity the lower the confining pressu re the stronger the liquefaction resistance. As a matter of fact, the difference here is whether absolute cyclic stress amplitude or cyclic stress ratio is used as loading intensity. Use of absolute stress amplitude to indicate liquefaction resistance may be appropriate in the study of static loading induced liquefaction, cyclic stress ratio is nevertheless more realistic when earthquake induced liquefaction is of concern. Y.P Vaid, J.D. Stedman and S.Sivathayalan (2001) showed that in cyclic loading the e ffect of increasing confining stress at a given static shear generally decreased the resistance to liquefaction. However, at the loosest states the increase in confining stress had little effect. B.1.10 Cyclic Stress Amplitude and Number of Cyclic Stres s Cycles In their laboratory study, Seed and Lee (1966) concluded that the larger the stress or stain, the lower the number of cycles required to induce liquefaction. Also the more the number of stress cycles to which the sand is subjected the more likely the liquefaction failure will occur. These two factors are directly related to the magnitude of cyclic loading. The effect of earthquake magnitude on liquefaction resistance of soils is apparent based on concept of cumulative damage proposed by Miner (1945 damage concept in soil liquefaction analysis was confirmed in studies concerning the validity of equivalent uniform stress cycle concept (Seed et al., 1975; Annaki and lee, Lee, 1997).

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343 B.1.11 Particle Size and Gradation Studies conducted by several researchers (Leed and Fitton, 1969; Wong et al., 1975; Ishihara et al., 1978) suggested that cyclic strength is the lowest with mean grain size, D 50 near 0.1 mm. increase or decrease in D 50 from 0.1mm tends to inc rease cyclic strength. Wang et al. (2002) also found that contrary to their expectation, well graded material was somewhat weaker than uniformly graded material. This unexpected observation was attributed to possible higher densification tendency and small er membrane penetration effect in well graded material which favored pore pressure generation. B.1.12 Pre straining Fanner et al. (2003) found that once a specimen has liquefied and reconsolidated to a denser structure, despite this densification, the specimen is much weaker to reliquefaction. Similar observation was also reported by Lee and Focht (1975). Study conducted by Mori et al. (1977) showed that specimens with prestraining by applying several loading cycles without causing liquefaction then rel easing excess pore pressure for consolidation exhibited stronger cyclic strength than those specimen without prestraining. B.1.13 Lateral Earth Pressure (K 0 ) and Over consolidation Ratio A Study on dense sand by Lee and Focht (1999) indicated an incr ease in cyclic stress ratio of about 30% for an OCR of 3. Ishihara et al. (1978) showed that cyclic strength increased as OCR and fines content increased. For specimens with no fines, a strength increase of 30% was observed for an increase in OCR from 1 to 2, while for the same OCR increase an 80% increase in cyclic strength was observed for specimens with 100% fines. Similar results produced from cyclic simple tests were reported by Seed and Peacock (1971).

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344 R. Segaldo et al. (1999) showed that the effect of K o on cyclic resistance of clean, uncemented, normally consolidated sand with D R of 30 95% can reasonably be taken into account by normalization with respect to the mean consolidation effective stress. When a change in K o is associated with overconsolidation, there is an additional increase in cyclic resistance that is probably due to a prestraining effect on the fabric or grain structure of the sand. The experimental data suggest that this additional increase in cyclic resistance ranges from about 10 40% at an OCR of 2 to about 25 100% at an OCR of 4. This range in the data may be par tly due to differences in soil, testing equipment, or stress path during consolidation and cyclic loading, indicating further research is necessary to quantify this effect more accurately. B.1.14 Consolidation Ratio, K c To simulate stress condition in a n embankment, anisotropic consolidation of specimen is required (Seed et al. 1975). In their earlier study regarding level ground liquefaction, Seed and Peacock (1970) pointed out that cyclic traixial test can produce desired stress changes only by consoli dating the specimen initially under isotropic condition. Under this condition, a constant normal stress and a controlled and continuously changing shear stress may be imposed along a 45 degree plane in the specimen. If any other consolidation pressure is u sed, there will be no plane in the specimen which will receive desired symmetrical changes in shear stress. In case of initially anisotropic stress condition, cyclic simple shear test can better simulate one dimensional cyclic loading condition. However, s tress variations due to earthquake can be very complicated in an embankment. One dimensional simplification is not appropriate and no proper test can be devised unless stress variations during earthquake can be realistically simulated.

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345 Castro and Poulos ( 1999) found that samples consolidated under higher K c would require a smaller increment in stresses to cause liquefaction, because at a higher K c , the specimen is closer to failure.

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346 APPENDIX C C. Procedure for Fabricating Rubber Membranes C .1 Initial Step C .1.1 Regulate the oven to 158±9° C (70±5°C). C .1.2 Be sure the proper amounts of coagulant and latex are in their respective containers (Always be sure there is enough coagulant and latex in t he containers to cover the largest mandrel, but not so much as to cause overflow during the dipping process.) 3.1.3 Stir the latex vigorously (Stir the latex and allow it to settle for 45 minutes before dipping the mandrel into it so that any bubbles caus ed by the stirring action will be dispersed. The latex must be stirred well before each use because a thick film of concentrated latex will form on the top. Inadequate stirring will cause non uniformity in the membrane thickness) C .2 Clean the mandrel(s) to be dipped. C .2.1 Wash mandrels with detergent. C .2.2 Rinse mandrels thoroughly in warm water. C .2.3 Place mandrels in the oven to dry. A 15 20 minute drying time is usually sufficient. C .3 Dip the mandrel in the coagulant while the mandrel is still warm from the drying process. C .3.1 Immerse the entire mandrel briefly into the coagulant. C .3.2 Allow excess coagulant to drip off the mandrel. C .3.3 Inspect the mandrel for unwetted spots. C .3.4 Place the mandrel back into the oven for 25 minutes. (T his step drives off the methanol and leaves a sticky film of calcium nitrate on the surface of the mandrel. An oven curing time of more than 25 minutes will cause the calcium nitrate to crystallize, resulting in spotty concentrations of latex on the mandre l)

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347 C .4 Dip the mandrel in the latex immediately after removing the mandrel from the oven. C .4.1 Prior to dipping the mandrel in the latex, inspect the surface of the latex in the container to make sure it is free from air bubbles and impurities. A spoon c an be used to scoop the surface clean. C .4.2 Slowly immerse the mandrel into the latex. Take care not to trap air between the latex and the mandrel surface. C .4.3 Dwell time begins when the mandrel is completely submerged. (A dwell time of ten seconds is sufficient to obtain a thin membrane with good strength and sensitivity. This is the type normally used by the Geotechnical Engineering Bureau. However, because membrane thickness is directly proportional to dwell time, a thinner or thicker membrane may b e obtained by decreasing or increasing the dwell time) C .4.4 Slowly remove the mandrel from the latex and allow the excess to drip off. C .4.5 Inspect the latex coating on the mandrel for any uncovered areas. (A very small hole can be repaired by gluing a small patch of rubber membrane over the hole with rubber cement. Do this after the membrane has oven cured and while the membrane is still on the mandrel. A patch with rounded corners is most effective. A large uncovered area is difficult to repair. The pr ocedure should be restarted from beginning) C .5 Initial curing stage. C .5.1 Place the mandrel in the oven at 158±9° F (70±5°C) for three hours. C .6 Strip the membrane off the mandrel. C .6.1 With the membrane still on the mandrel, cut along the top and bo ttom edges of the membrane with a sharp razor blade or Exact o knife. The cut must be smooth and even. Apply a thin coat of rubber cement 0.5 in. (12.7 mm) wide along each edge. Allow the

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348 cement to become tacky, and then carefully roll the membrane edges d own 0.5 in. (12.7 mm). This procedure creates strong, tear resistant edges. C .6.2 Dust the membrane with talcum powder. This prevents the membrane from sticking to itself when being stripped from the mandrel. C .6.3 Carefully pull the membrane down the ma ndrel, stopping at intervals to dust the inside portion of the membrane with talcum powder. (Avoid excessively stretching the membrane as it is not fully cured and will tear easily) C .7 Final curing stages C .7.1 Completely submerge the membrane in warm water for three hours or in cold water overnight. This will remove any latent ammonia from the membrane. C .7.2 Remove the membrane from the water and allow it to air dry. (Do not subject the membrane to stretc hing until it is completely dry. It is very weak and will tear quite easily) C .7.3 Trim any rough edges from the membrane. C .7.4 Store the membrane in a dry place away from any light source. (Petroleum products will destroy natural rubber. Therefore, do n ot expose the membranes to petroleum based oils, petroleum jelly, etc.) Tight closing cardboard boxes of sufficient size would be a good way to store the membranes.

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349 APPENDIX D D. Cyclic Triaxial and Hollow Cylinder Tests Results Cyclic Triaxial Test Relative Density after saturation (%) Stress Ratio Void Ratio No. cycles to reach initial liquefaction Deviator Stress (psi) Consolidation Pressure (psi) Fine Content (%) Plasticity Index, PI Sample No. 30.22 0.2 0.7315 32 6 15 5 20 1 30.32 0.3 0.7312 19 9 15 5 20 2 30.01 0.4 0.7320 10 12 15 5 20 3 30.01 0.2 0.7320 112 12 30 5 20 4 29.8 0.3 0.7325 63 18 30 5 20 5 30.11 0.4 0.7317 30 24 30 5 20 6 31.75 0.2 0.7278 24 6 15 10 20 7 30.67 0.3 0.7304 10 9 15 10 20 8 29.95 0.4 0.7321 5 12 15 10 20 9 30.16 0.2 0.7316 48 12 30 10 20 10 30.4 0.3 0.7310 30 18 30 10 20 11 30.34 0.4 0.7312 18 24 30 10 20 12 30.46 0.2 0.7309 10 6 15 15 20 13 30.16 0.3 0.7316 5 9 15 15 20 14 30.34 0.4 0.7312 3 12 15 15 20 15 30.1 0.2 0.7318 35 12 30 15 20 16 29.92 0.3 0.7322 22 18 30 15 20 17 29.62 0.4 0.7329 13 24 30 15 20 18 29.92 0.2 0.7322 28 6 15 25 20 19 29.3 0.3 0.7337 14 9 15 25 20 20 29.76 0.4 0.7326 8 12 15 25 20 21 28.64 0.2 0.7353 74 12 30 25 20 22 29.76 0.3 0.7326 36 18 30 25 20 23 29.44 0.4 0.7333 21 24 30 25 20 24 29.76 0.2 0.7326 45 6 15 35 20 25 29.94 0.3 0.7321 26 9 15 35 20 26 29.88 0.4 0.7323 14 12 15 35 20 27 30.12 0.2 0.7317 128 12 30 35 20 28 30.6 0.3 0.7306 72 18 30 35 20 29 30.24 0.4 0.7314 32 24 30 35 20 30 29.64 0.2 0.7329 56 6 15 45 20 31 29.19 0.3 0.7339 32 9 15 45 20 32 29.29 0.4 0.7337 15 12 15 45 20 33 29.74 0.2 0.7326 153 12 30 45 20 34 30.34 0.3 0.7312 86 18 30 45 20 35

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350 29.92 0.4 0.7322 38 24 30 45 20 36 44.68 0.2 0.6968 118 6 15 5 20 37 45.39 0.3 0.6951 65 9 15 5 20 38 44.99 0.4 0.6960 35 12 15 5 20 39 45.45 0.2 0.6949 256 12 30 5 20 40 45.15 0.3 0.6956 115 18 30 5 20 41 45.27 0.4 0.6954 50 24 30 5 20 42 45.03 0.2 0.6959 78 6 15 10 20 43 44.49 0.3 0.6972 36 9 15 10 20 44 44.72 0.4 0.6967 18 12 15 10 20 45 44.96 0.2 0.6961 104 12 30 10 20 46 45.11 0.3 0.6957 51 18 30 10 20 47 44.8 0.4 0.6965 28 24 30 10 20 48 45.21 0.2 0.6955 36 6 15 15 20 49 44.93 0.3 0.6962 24 9 15 15 20 50 45.27 0.4 0.6954 14 12 15 15 20 51 44.84 0.2 0.6964 51 12 30 15 20 52 44.41 0.3 0.6974 28 18 30 15 20 53 44.72 0.4 0.6967 15 24 30 15 20 54 45.09 0.2 0.6958 86 6 15 25 20 55 45.17 0.3 0.6956 40 9 15 25 20 56 44.7 0.4 0.6967 20 12 15 25 20 57 45.01 0.2 0.6960 125 12 30 25 20 58 44.92 0.3 0.6962 63 18 30 25 20 59 45.45 0.4 0.6949 32 24 30 25 20 60 45.21 0.2 0.6955 102 6 15 35 20 61 44.92 0.3 0.6962 50 9 15 35 20 62 44.92 0.4 0.6962 26 12 15 35 20 63 44.96 0.2 0.6961 215 12 30 35 20 64 45.03 0.3 0.6959 104 18 30 35 20 65 44.8 0.4 0.6965 52 24 30 35 20 66 60.09 0.2 0.6598 145 6 15 5 20 67 60.2 0.3 0.6595 86 9 15 5 20 68 59.97 0.4 0.6601 55 12 15 5 20 69 59.92 0.2 0.6602 328 12 30 5 20 70 60.17 0.3 0.6596 124 18 30 5 20 71 60.01 0.4 0.6600 90 24 30 5 20 72 59.82 0.2 0.6604 104 6 15 10 20 73 59.71 0.3 0.6607 56 9 15 10 20 74 60.09 0.4 0.6598 34 12 15 10 20 75 59.86 0.2 0.6603 213 12 30 10 20 76 59.63 0.3 0.6609 95 18 30 10 20 77 59.55 0.4 0.6611 75 24 30 10 20 78 59.67 0.2 0.6608 55 6 15 15 20 79 60.4 0.3 0.6590 40 9 15 15 20 80

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351 60.47 0.4 0.6589 28 12 15 15 20 81 60.03 0.2 0.6599 98 12 30 15 20 82 60.2 0.3 0.6595 78 18 30 15 20 83 59.94 0.4 0.6601 51 24 30 15 20 84 60.15 0.2 0.6596 112 6 15 25 20 85 59.92 0.3 0.6602 60 9 15 25 20 86 60.2 0.4 0.6595 41 12 15 25 20 87 59.86 0.2 0.6603 246 12 30 25 20 88 60.15 0.3 0.6596 97 18 30 25 20 89 60.2 0.4 0.6595 81 24 30 25 20 90 59.94 0.2 0.6601 126 6 15 35 20 91 59.71 0.3 0.6607 64 9 15 35 20 92 60.09 0.4 0.6598 48 12 15 35 20 93 60.3 0.2 0.6593 289 12 30 35 20 94 60.2 0.3 0.6595 107 18 30 35 20 95 60.15 0.4 0.6596 88 24 30 35 20 96

PAGE 380

352 Cyclic Hollow Cylinder Test Relative Density after saturation (%) Stress ratio Void Ratio No. cycles to reach initial liquefaction Deviator Stress (psi) consolidation pressure (psi) Fine Content (%) Plasticity index, PI Sample No. 30.16 0.4 0.7316 8 6 15 5 20 1 29.08 0.4 0.7342 28 12 30 5 20 2 30.39 0.3 0.7311 7 4.5 15 10 20 3 30.21 0.3 0.7315 26 9 30 10 20 4 29.44 0.2 0.7333 7 3 15 15 20 5 29.8 0.2 0.7325 26 6 30 15 20 6 30.6 0.3 0.7306 7 4.5 15 25 20 7 31.73 0.3 0.7278 26 9 30 25 20 8 29.73 0.4 0.7326 8 6 15 35 20 9 29.7 0.4 0.7327 27 12 30 35 20 10 60.21 0.4 0.6595 19 6 15 5 20 11 61.43 0.4 0.6566 40 12 30 5 20 12 61.9 0.3 0.6554 18 4.5 15 10 20 13 61.01 0.3 0.6576 39 9 30 10 20 14 61.9 0.2 0.6554 18 3 15 15 20 15 59.77 0.2 0.6606 38 6 30 15 20 16 62.32 0.3 0.6544 19 4.5 15 25 20 17 58.94 0.3 0.6625 38 9 30 25 20 18 60.59 0.4 0.6586 18 6 15 35 20 19 59.71 0.4 0.6607 39 12 30 35 20 20

PAGE 381

353 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D1. Test results (a) and (b) on Sample No.1. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. of Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Excess Pore Water Pressure (psi) No. of Cycles to Liquefaction

PAGE 382

354 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D2. Test results (a) and (b) on Sample No.2. -10 -8 -6 -4 -2 0 2 4 6 8 10 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No. of Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 Excess Pore Water Pressure (psi) No. of Cycles to Liquefaction

PAGE 383

355 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefac tion Figure D3. Test results (a) and (b) on Sample No.3. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 16 18 Cyclic Deviator Stress (psi) No. of Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 20 0 2 4 6 8 10 12 14 Excess Pore Water Pressure (psi) No. of Cycles to Liquefaction

PAGE 384

356 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D4. Test results (a) and (b) on Sample No.4. -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 385

357 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D5. Test results (a) and (b) on Sample No.5. -20 -15 -10 -5 0 5 10 15 20 25 0 10 20 30 40 50 60 70 80 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 386

358 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D6. Test results (a) and (b) on Sample No.6. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 387

359 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cyc les to liquefaction Figure D7. Test results (a) and (b) on Sample No.7. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 388

360 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D8. Test results (a) and (b) on Sample No.8. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 16 18 Cyclic Deviator Stress (psi) No. Cyclic to Liquefaction 0 2 4 6 8 10 12 14 16 18 20 0 2 4 6 8 10 12 14 16 Excess Pore Pressure (psi) No. Cyclic to Liquefaction

PAGE 389

361 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D9. Test results (a) and (b) on Sample No.9. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 Cyclic Deviator Stress (psi) No. Cyclic to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 1 2 3 4 5 6 7 8 9 10 Excess Pore Pressure (psi) No.Cyclic to Liquefaction

PAGE 390

362 (a) Cyclic deviator stress(psi) versus number of cy cles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D10. Test results (a) and (b) on Sample No.10. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 Cyclic Deviator Stress (psi) No. Cyclic to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 Excess Pore Water Pressure (psi) No. Cycllic to liquefaction

PAGE 391

363 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus numbe r of cycles to liquefaction Figure D11. Test results (a) and (b) on Sample No.11. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. Cyclic to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 50 Excess Pore Water Pressure (psi) No. Cycllic to liquefaction

PAGE 392

364 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D12. Test results (a) and (b) on Sample No.12. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No. Cyclic to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 Excess Pore Water Pressure (psi) No.Cyclic to Liquefaction

PAGE 393

365 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D13. Test results (a) and (b) on Sample No.13. -8 -6 -4 -2 0 2 4 6 8 0 2 4 6 8 10 12 14 16 Cyclic Deviator stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 20 0 2 4 6 8 10 12 14 Excess Pore Pressure (psi) No. Cycles to liquefaction

PAGE 394

366 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D14. Test results (a) and (b) on Sample No.14. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 Cyclic Deviator Stress (psi) No. Cycless to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 1 2 3 4 5 6 7 8 9 10 Excess Pore Pressure (psi) No.Cyclic to Liquefaction

PAGE 395

367 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D15. Test results (a) and (b) on Sample No.15. -15 -10 -5 0 5 10 15 0 1 2 3 4 5 6 7 8 9 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 Ecess pore pressure (psi) No.Cycles to Liquefaction

PAGE 396

368 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D16. Test results (a) and (b) on Sample No.16. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 50 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 397

369 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D17. Test results (a) and (b) on Sample No.17. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 398

370 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D18. Test results (a) and (b) on Sample No.18. -30 -20 -10 0 10 20 30 0 2 4 6 8 10 12 14 16 18 20 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 Excess pore pressure (psi) No.Cycles to Liquefaction

PAGE 399

371 (a) Cyclic deviator stress(psi) versus number of cycles to li quefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D19. Test results (a) and (b) on Sample No.19. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 40 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No, Cycles to liquefaction

PAGE 400

372 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 20. Test results (a) and (b) on Sample No.20. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 16 18 20 Cyclic Deviator Streess(psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 2 4 6 8 10 12 14 16 18 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 401

373 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D 21. Test results (a) and (b) on Sample No.21. -10 -8 -6 -4 -2 0 2 4 6 8 10 0 2 4 6 8 10 12 14 Cyclic Deviator pressure (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 14 Excess pore pressture (psi) No. cycles to liquefaction

PAGE 402

374 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 22. Test results (a) and (b) on Sample No.22 . -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 80 90 100 Cyclic Deviator Stress (psi) No cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 403

375 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 23. Test results (a) and (b) on Sample No.23. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 5 10 15 20 25 30 35 40 45 Cyclic Deviator stress (psi) No.Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess Pore pressure(psi) No.Cycles to liquefaction

PAGE 404

376 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 24. Test results (a) and (b) on Sample No.24. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) NO. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 405

377 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of c ycles to liquefaction Figure D 25. Test results (a) and (b) on Sample No.25. -8 -6 -4 -2 0 2 4 6 8 0 10 20 30 40 50 60 Cyclic Deviator stress (psi) no.cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 40 45 50 Excess pore pressuree (psi) No. Cycles to liquefaction

PAGE 406

378 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 26. Test results (a) and (b) on Sample No.26. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 Excess pore pressure (psi) No,Cycles to liquefaction

PAGE 407

379 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 27. Test results (a) and (b) on Sample No.27. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 16 18 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 2 4 6 8 10 12 14 16 18 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 408

380 (a) Cyclic deviator stress(psi) versus numb er of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 28. Test results (a) and (b) on Sample No.28. -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 140 160 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 Excess pore pressure (psi) No. Cycles to liquefion

PAGE 409

381 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 29. Test results (a) and (b) on Sample No.29. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 10 20 30 40 50 60 70 80 90 100 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 410

382 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D 30. Test results (a) and (b) on Sample No.30. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess Pore Pressure (psi) No. Cycles to liquefaction

PAGE 411

383 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 31. Test results (a) and (b) on Sample No.31. -8 -6 -4 -2 0 2 4 6 8 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 Exces pore pressure (psi) No. cycles to liquefaction

PAGE 412

384 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 32. Test results (a) and (b) on Sample No.32. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 Cyclic Deviator Sress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Excess Pore Pressure (psi) No. Cycles to Liquefaction

PAGE 413

385 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 33. Test results (a) and (b) on Sample No.33. -15 -10 -5 0 5 10 15 0 2 4 6 8 10 12 14 16 18 20 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 2 4 6 8 10 12 14 16 Excess pore pressure (psi0 No. Cycles to liquefaction

PAGE 414

386 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of c ycles to liquefaction Figure D 34. Test results (a) and (b) on Sample No.34. -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 140 160 180 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 160 180 Excess pore pressure (psi) No. cycles to lqiuefaction

PAGE 415

387 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 35. Test results (a) and (b) on Sample No.35. -20 -15 -10 -5 0 5 10 15 20 25 0 10 20 30 40 50 60 70 80 90 100 Cyclic Deviator Stress(psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 416

388 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 36. Test results (a) and (b) on Sample No.36. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 45 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 417

389 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 37. Test results (a) and (b) on Sample No.37. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 140 160 Cyclic Deviator Stress (psi) No. cycle of Liueqfaction 0 2 4 6 8 10 12 14 16 0 20 40 60 80 100 120 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 418

390 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 38. Test results (a) and (b) on Sample No.38. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 80 Cyclic Deviator Stress(psi) No. cycles to liueqfaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 70 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 419

391 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 39. Test results (a) and (b) on Sample No.39. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 Cyclic Deviator Stress (psi) No. cycles to liiquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Excess Pore water pressure (psi) No. Cycles to liquefaction

PAGE 420

392 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D 40. Test results (a) and (b) on Sample No.40. -15 -10 -5 0 5 10 15 0 50 100 150 200 250 300 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 300 Excess pore pressure (psi) No. cycles to lqiuefaction

PAGE 421

393 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 41. Test results (a) and (b) on Sample No.41. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 20 40 60 80 100 120 140 Cyclic Deviator Stress(psi) No.Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction

PAGE 422

394 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 42. Test results (a) and (b) on Sample No.42. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 Cyclic Deviator stress (psi) No. cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 423

395 (a) Cyclic deviator stress(psi) versus number of cycles to liqu efaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 43. Test results (a) and (b) on Sample No.43. -8 -6 -4 -2 0 2 4 6 8 0 10 20 30 40 50 60 70 80 90 Cyclic Deviator stress (psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 10 20 30 40 50 60 70 80 90 excess pore pressure (psi) No. cycles to liquefaction

PAGE 424

396 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 44. Test results (a) and (b) on Sample No.44. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 50 Cyclic Deviator Stress(psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 425

397 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 45. Test results (a) and (b) on Sample No.45. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 excess pore pressure (psi) no. cycles to lqiuefaction

PAGE 426

398 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 46 . Test results (a) and (b) on Sample No.4 6 . -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 427

399 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D 47. Test results (a) and (b) on Sample No.47. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 Excess pore pressure (psi) No. cycles to lqiuefaction

PAGE 428

400 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 48. Test results (a) and (b) on Sample No.48. -40 -30 -20 -10 0 10 20 30 40 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No.Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 429

401 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 49. Test results (a) and (b) on Sample No.49. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 40 45 50 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 430

402 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 50. Test results (a) and (b) on Sample No.50 . -40 -30 -20 -10 0 10 20 30 40 0 5 10 15 20 25 30 35 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 30 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 431

4 03 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 51. Test results (a) and (b) on Sample No.51. -40 -30 -20 -10 0 10 20 30 40 0 5 10 15 20 25 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 2 4 6 8 10 12 14 16 18 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 432

404 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefac tion Figure D 52. Test results (a) and (b) on Sample No.52. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 433

405 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 53. Test results (a) and (b) on Sample No.53. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No.Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 434

406 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 54. Test results (a) and (b) on Sample No.54. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 5 10 15 20 25 30 35 0 2 4 6 8 10 12 14 16 18 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 435

407 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 55. Test results (a) and (b) on Sample No.55. -6 -4 -2 0 2 4 6 0 20 40 60 80 100 120 Cyclic Deviator Stress (psi) No.Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 436

408 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (a) Excess Pore water pressure versus number of cyc les to liquefaction Figure D 56. Test results (a) and (b) on Sample No.56. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 40 45 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 437

409 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 57. Test results (a) and (b) on Sample No.57. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 438

410 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 58. Test results (a) and (b) on Sample No.58. -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 439

411 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 59. Test results (a) and (b) on Sample No.59. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 80 Cyclic Deviator Stres (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 440

412 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus num ber of cycles to liquefaction Figure D 60. Test results (a) and (b) on Sample No.60. -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 45 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No.cycles to liquefaction

PAGE 441

413 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 61. Test results (a) and (b) on Sample No.61. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 140 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 442

414 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 62. Test results (a) and (b) on Sample No.62. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 443

415 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 63. Test results (a) and (b) on Sample No.63. -40 -30 -20 -10 0 10 20 30 40 0 5 10 15 20 25 30 35 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 444

416 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressur e versus number of cycles to liquefaction Figure D 64. Test results (a) and (b) on Sample No.64. -15 -10 -5 0 5 10 15 0 50 100 150 200 250 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 445

417 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 65. Test results (a) and (b) on Sample No.65. -25 -20 -15 -10 -5 0 5 10 15 20 25 0 20 40 60 80 100 120 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 446

418 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 66. Test results (a) and (b) on Sample No.66. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 10 20 30 40 50 60 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 447

419 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 67. Test results (a) and (b) on Sample No.67. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 140 160 180 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 20 40 60 80 100 120 140 160 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 448

420 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 68. Test results (a) and (b) on Sample No.68. -10 -8 -6 -4 -2 0 2 4 6 8 10 0 10 20 30 40 50 60 70 80 90 100 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 449

421 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (a) Excess Pore water pressure versus number of cycles to liquefaction Figure D 69. Test results (a) and (b) on Sample No.69. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 10 20 30 40 50 60 excess pore pressure (psi) No. cycles to liquefaction

PAGE 450

422 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (a) Excess Pore water pressure versus number of cycles to liquefaction Figure D 70. Test results (a) and (b) on Sample No.70. -15 -10 -5 0 5 10 15 0 50 100 150 200 250 300 350 400 Cyclic Deviator stress (psi) No. cycles to liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 300 350 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 451

423 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 71. Test results (a) and (b) on Sample No.71. -20 -15 -10 -5 0 5 10 15 20 0 20 40 60 80 100 120 140 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 140 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 452

424 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 72. Test results (a) and (b) on Sample No.72. -30 -20 -10 0 10 20 30 0 20 40 60 80 100 120 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 453

425 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefa ction Figure D 73. Test results (a) and (b) on Sample No.73. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 Cyclic Deviator stress (psi) No.Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 20 40 60 80 100 120 Excess pore pressure (psi) No.cycles to liquefaction

PAGE 454

426 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 74. Test results (a) and (b) on Sample No.74. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 455

427 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 75 . Test results (a) and (b) on Sample No.75 . -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 456

428 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 76 . Test results (a) and (b) on Sample No.76 . -15 -10 -5 0 5 10 15 0 50 100 150 200 250 Cyclic Deviator Stress (psi) No. cycles to liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 457

429 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cyc les to liquefaction Figure D 7 7 . Test results (a) and (b) on Sample No.7 7 . -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 458

430 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 7 8 . Test results (a) and (b) on Sample No.7 8 . -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 80 90 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 10 20 30 40 50 60 70 80 90 Excess pore pressure (psi) No. cycles to lquefaction

PAGE 459

431 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 7 9 . Test results (a) and (b) on Sample No.79 . -8 -6 -4 -2 0 2 4 6 8 0 10 20 30 40 50 60 70 Cyclic Deviator stress (psi) No. cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 Excess pore pressure (psi) No.cycles to liquefaction

PAGE 460

432 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 80 . Test results (a) and (b) on Sample No.80 . -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 Cyclic deviator stress (psi) No. cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 40 45 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 461

433 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 81 . Test results (a) and (b) on Sample No.81 . -30 -20 -10 0 10 20 30 0 5 10 15 20 25 30 35 40 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 Excess pore pressure(psi) No. cycles to lqiuefaction

PAGE 462

434 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefac tion Figure D 8 2 . Test results (a) and (b) on Sample No.8 2 . -15 -10 -5 0 5 10 15 0 20 40 60 80 100 120 Cyclic DEviator stress (psi) No. cycles to liqufaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 463

435 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 8 3 . Test results (a) and (b) on Sample No.83 . -25 -20 -15 -10 -5 0 5 10 15 20 25 0 10 20 30 40 50 60 70 80 90 Cyclic DEviator stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 464

436 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 84. Test results (a) and (b) on Sample No.84. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 Cyclic Deviator Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 465

437 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 85. Test results (a) and (b) on Sample No.85. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 140 Cyclic Deviator stress (psi) No. cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 20 40 60 80 100 120 Excess pore pressure (psi) No. cycles to liquefacttion

PAGE 466

438 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cy cles to liquefaction Figure D 86. Test results (a) and (b) on Sample No.86. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 Cyclic Deviator Stress(psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 70 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 467

439 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 87. Test results (a) and (b) on Sample No.87. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 50 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 5 10 15 20 25 30 35 40 45 Excess pore pressure(psi) No. cycles to liquefaction

PAGE 468

440 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 88. Test results (a) and (b) on Sample No.88. -15 -10 -5 0 5 10 15 0 50 100 150 200 250 300 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 300 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 469

441 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 89. Test results (a) and (b) on Sample No.89. -20 -15 -10 -5 0 5 10 15 20 0 20 40 60 80 100 120 Cyclic deviator stress (psi) No. Cyclles to liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure (psi) No.cycles to liquefaction

PAGE 470

442 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 90. Test results (a) and (b) on Sample No.90. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 80 90 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 471

443 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefac tion Figure D 91. Test results (a) and (b) on Sample No.91. -8 -6 -4 -2 0 2 4 6 8 0 20 40 60 80 100 120 140 160 Cyclic Deviator stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 20 40 60 80 100 120 140 Excess pore pressure (psi) No. Cycles to liquefaction

PAGE 472

444 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 92. Test results (a) and (b) on Sample No.92. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 70 80 Cyclic Deviator stress (psi) No. cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 70 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 473

445 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 93. Test results (a) and (b) on Sample No.93. -15 -10 -5 0 5 10 15 0 10 20 30 40 50 60 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 10 20 30 40 50 60 Excess pore preesure (psi) No. cycles to liquefaction

PAGE 474

446 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 94. Test results (a) and (b) on Sample No.94. -15 -10 -5 0 5 10 15 0 50 100 150 200 250 300 350 Cyclic Deviator Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 50 100 150 200 250 300 350 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 475

447 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cyc les to liquefaction Figure D 95. Test results (a) and (b) on Sample No.95. -20 -15 -10 -5 0 5 10 15 20 0 20 40 60 80 100 120 140 Cyclic Deviator stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 20 40 60 80 100 120 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 476

448 (a) Cyclic deviator stress(psi) versus number of cycles to liquefaction (b) Excess Pore water pressure versus number of cycles to liquefaction Figure D 96. Test results (a) and (b) on S ample No.96. -30 -20 -10 0 10 20 30 0 10 20 30 40 50 60 70 80 90 100 Cyclic Deviator stress(psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 10 20 30 40 50 60 70 80 90 100 Excess pore pressure (psi) NO. cycles to liquefaction

PAGE 477

449 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 1. Test Results (a) and (b) on sample No.1 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 2 4 6 8 10 12 14 16 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess Pore Water Pressure (psi) No. Cycles to Liquefaction

PAGE 478

450 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 2. Test Results (a) and (b) on sample No.2 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 Cyclic Shear Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 Excess Pore Water (psi) No. Cycles to liquefaction

PAGE 479

451 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 3. Test Results (a) and (b) on Sample No.3 in CHCT. -6 -4 -2 0 2 4 6 0 2 4 6 8 10 12 Cyclic Shear Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore pressure (psi) No. Cycles to Liquefaction

PAGE 480

45 2 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liq uefaction Figure E 4. Test Results (a) and (b) on Sample No.4 in CHCT. -10 -8 -6 -4 -2 0 2 4 6 8 10 0 5 10 15 20 25 30 35 Cyclic Shear Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 481

453 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 5. Test Results (a) and (b) on Sample No.5 in CHCT. -4 -3 -2 -1 0 1 2 3 4 0 2 4 6 8 10 12 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Cyclic Shear Stress (psi) No. Cycles to liquefaction

PAGE 482

454 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 6. Test Results (a) and (b) on Sample No.6 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 483

455 (a) Cyclic shear stress (psi) versus number o f cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 7. Test Results (a) and (b) on Sample No.7 in CHCT. -6 -4 -2 0 2 4 6 0 2 4 6 8 10 12 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 2 4 6 8 10 12 14 16 18 0 2 4 6 8 10 12 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 484

456 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 8. Test Results (a) and (b) on Sample No.8 in CHCT. -12 -10 -8 -6 -4 -2 0 2 4 6 8 10 0 5 10 15 20 25 30 35 Cyclic Shear Strength (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 485

457 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 9. Test Results (a) and (b) on Sample No.9 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 2 4 6 8 10 12 Cyclic Shear Stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 0 1 2 3 4 5 6 7 8 9 10 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 486

458 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 10. Test Results (a) and (b) on Sample No.10 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 Cyclic Shear Stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 487

459 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to l iquefaction Figure E 11. Test Results (a) and (b) on Sample No.11 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 Cyclic shear stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 488

460 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 12. Test Results (a) and (b) on Sample No.12 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 50 Cyclic shear stress (psi) No. cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 50 Excess pore pressure (psi) No. Cycles to Liquefaction

PAGE 489

461 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 13. Test Results (a) and (b) on Sample No.13 in CHCT. -6 -4 -2 0 2 4 6 0 5 10 15 20 25 Cyclic Shear stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 excess pore pressure (psi) No. cycles to liquefaction

PAGE 490

462 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 14. Test Results (a) and (b) on Sample No.14 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 50 Cyclic shear stress (psi) No. cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 50 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 491

463 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 15. Test Results (a) and (b) on Sample No.15 in CHCT. -4 -3 -2 -1 0 1 2 3 4 0 5 10 15 20 25 30 Cyclic Shear Stress(psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 Excess pore poressure (psi) No. Cycles to liquefaction

PAGE 492

464 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 16. Test Results (a) and (b) on Sample No.16 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 35 40 45 50 Cyclic shear stress (psi) No. Cycles to liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 50 Excess pore pressure (psi) No. cycles to lqiuefaction

PAGE 493

465 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 17. Test Results (a) and (b) on Sample No.17 i n CHCT. -6 -4 -2 0 2 4 6 0 5 10 15 20 25 Cyclic Shear stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 494

466 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 18. Test Results (a) and (b) on Sample No.18 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 Excess pore pressure (psi) No. cycles to liquefaction

PAGE 495

467 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 19. Test Results (a) and (b) on Sample No.19 in CHCT. -8 -6 -4 -2 0 2 4 6 8 0 5 10 15 20 25 30 Cyclic shear stress (psi) No. Cycles to liquefaction 0 2 4 6 8 10 12 14 16 18 0 5 10 15 20 25 excess pore pressure (psi) No. cycles to liquefaction

PAGE 496

468 (a) Cyclic shear stress (psi) versus number of cycles to liquefaction. (b) Excess Pore water pressure versus number of cycles to liquefaction Figure E 20. Test Results (a) and (b) on Sample No.20 in CHCT. -15 -10 -5 0 5 10 15 0 5 10 15 20 25 30 35 40 45 Cyclic Shear Stress (psi) No. Cycles to Liquefaction 0 5 10 15 20 25 30 35 0 5 10 15 20 25 30 35 40 45 50 Excess pore pressure (psi) No. cycless to liquefaction