the increase in shaft capacity with time for friction

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LSU Historical Dissertations and Theses Graduate School

1996

The Increase in Shaft Capacity With Time forFriction Piles Driven Into Saturated Clay.Hani Hasan TitiLouisiana State University and Agricultural & Mechanical College

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THE INCREASE IN SHAFT CAPACITY W ITH TIME FOR FRICTION PILES DRIVEN INTO SATURATED CLAY

A Dissertation

Submitted to the Graduate Faculty of the Louisiana State University and

Agricultural and Mechanical College in partial fulfillment of the

requirements for the degree of Doctor of Philosophy

m

The Department of Civil and Environmental Engineering

byHani Hasan Titi

B.S.C.E., Yarmouk University, Irbid, Jordan, 1987 M.S.C.E., Jordan University of Science and Technology, Irbid, Jordan, 1990

December 1996


UMI Number: 9712861

UMI Microform 9712861 Copyright 1997, by UMI Company. All rights reserved.

This microform edition is protected against unauthorized copying under Title 17, United States Code.

UMI300 North Zeeb Road Ann Arbor, MI 48103


Dedicated

To M y Parents

ii


ACKNOWLEDGEMENTS

I would like to express my deepest appreciation to my advisor Professor G. Wije

Wathugala for his friendship, guidance, and support throughout this research, project.

His comments and suggestions have been invaluable.

My sincere gratitude is extended to Professor George Z. Voyiadjis. His help

and support during my study is acknowledged. I would like to thank Professor

Mehmet T. Tumay. His help, valuable comments, and suggestions are acknowledged

with gratitude. My thanks also goes to Professor Roger K. Seals for his valuable

comments. I would like to thank Professor Bruce Sbarky for his serving in the

examination committee.

I would like here to remember the memory of Professor Yalcin B. Acar who was

a committee member. His friendship will always remain in my thoughts.

This research performed herein was partially supported by Louisiana Transporta

tion Research Center (LTRC) and by the National Science Foundation (NSF) through

the grant No. CMS-9410482. These supports are gratefully acknowledged.

My t h a n k s extended to Dr. Akram Alshawabkeh and to Mr. Surajit Pal for

their friendship and support. I would like also to thank Mr. Manaa’ Mesleh for his

friendship and support. The help of my friend Mr. Ibrahim Mansi is acknowledged

with appreciation. I would like to t h ank my parents and siblings for continuous

support and encouragements.

To my wife Eman and my daughter Miriam I owe the most. This work would

have not been possible without their support and love.

iii


CONTENTS

D E D IC A T IO N ...................................................................................................... ii

A C K N O W L E D G E M E N T S............................................................................. iii

L IS T O F T A B L E S ..................................................................................................... vii

LIST OF F IG U R E S ................................................................................................ viii

ABSTRACT .............................................................. xvii

CHAPTER

1 INTRODUCTION ....................................................................................... 11.1 B ack g ro u n d .......................................................t .......................................... 11.2 Objectives and Scope of the R esearch....................................................... 31.3 Organization of the M an u sc rip t................................................................. 4

2 BACKGROUND .......................................................................................... 62.1 In troduction .................................................................................................... 62.2 Effects of Pile Driving on the Clay-Water S y s te m ................................ 7

2.2.1 Disturbance Effects on Soil P roperties .......................................... 82.2.2 Changes in Pore Water P re s s u re s ................................................. 82.2.3 Equalization of Pore Water Pressures and S tresses ................... 11

2.3 Pile S e tu p ........................................................................................................ 122.4 Soil Sensitivity .............................................................................................. 182.5 Methods of Predicting Pile Shaft C a p a c ity ............................................. 192.6 Pile-Setup Related S tu d ies .......................................................................... 21

3 HIERARCHICAL SINGLE SURFACE M ODELING APPROACH 263.1 In troduction .................................................................................................... 263.2 The 5*-Series of HiSS Constitutive M o d e ls ............................................. 27

3.2.1 Basics of the Incremental Theory of P la s t ic i ty .......................... 273.2.2 Yield Surface, F ................................................................................. 283.2.3 Hardening Function, O p , ................................................................. 323.2.4 Interpolation Functions.................................................................... 343.2.5 Potential Function, Q ....................................................................... 36

3.3 The Proposed Interface HiSS-^,- M odel.................................................... 363.3.1 Potential Function, Q ....................................................................... 373.3.2 Translation R u le ................................................................................. 39

3.4 Evaluation of the M o d e l ............................................................................. 443.4.1 Field Investigation on Sabine C la y ................................................. 44

iv


3.4.2 Material Parameters ........................................................................ 453.4.3 Implementation of the Model into Computer Procedures . . . 483.4.4 Capabilities and Properties of the M o d e l .................................... 483.4.5 Calibration of the M o d e l ................................................................. 51

4 S IM U L A T IO N O F P IL E IN S T A L L A T IO N ............................................ 604.1 In troduction ..................................................................................................... 604.2 Methods of Pile Driving S im ulation ........................................................... 614.3 Strain Path M ethod........................................................................................ 63

4.3.1 Limitations ........................................................................................ 634.3.2 Closed-Ended P i l e s ........................................................................... 634.3.3 Open-Ended P ile s .............................................................................. 67

4.3.3.1 Ring S o u rc e ........................................................................ 704.3.3.2 Superposing a Uniform Steady Flow on the Source-

Ring F lo w ........................................................................... 714.3.3.3 Strain F i e l d ....................................................................... 73

4.4 Numerical Simulation of Pile D riv in g ........................................................ 78

5 V E R IF IC A T IO N : P IL E D R IV IN G , S U B S E Q U E N T C O N SO L ID A T IO N A N D P IL E LO AD T E S T S .............................................................. 845.1 In troduction ..................................................................................................... 845.2 Field Tests at Sabine Site ........................................................................... 85

5.2.1 Test P rocedures................................................................................. 855.2.2 Pile Segment M o d e ls ....................................................................... 885.2.3 Results of Field Experim ents.......................................................... 88

5.3 The Finite Element Program A B A Q U S/S tandard................................ 925.3.1 Analysis of Porous M ed ia ................................................................ 925.3.2 Implementation of the HiSS-£Jt- Model in A B A Q U S ................. 955.3.3 Finite Element M e s h ....................................................................... 96

5.4 Pile D riv in g ........................................................................................................1025.5 Consolidation and Pile Load T e s ts .................................................................1085.6 Pile S e tu p ...........................................................................................................135

6 N U M E R IC A L SIM U L A T IO N O F T H E B E H A V IO R O F C L O S E D -E N D E D P IL E S .....................................................................................................1376.1 In troduction ....................................................................................................... 1376.2 Field Experim ents..............................................................................................1386.3 Finite Element M e s h ....................................................................................... 1386.4 Pile D r iv in g ....................................................................................................... 1436.5 Consolidation and Pile Load T e s ts ................................................................ 1486.6 Pile S e tu p ...........................................................................................................167

7 N U M E R IC A L S IM U L A T IO N O F T H E B E H A V IO R O F O P E N -E N D E D P IL E S .....................................................................................................1697.1 In troduction ....................................................................................................... 169

v


7.2 Field Experim ents............................................................................................. 1707.3 Finite Element M esh ........................................................................................1707.4 Pile D r iv in g .......................................................................................................178

7.4.1 The 3 in. x 0.125 in. Open-ended Pile Segment Model . . . . 1797.4.2 The 3 in. x 0.065 in. Open-ended Pile Segment Model . . . . 179

7.5 Consolidation and Pile Load T e s ts ................................................................ 1847.5.1 The 3 in. x 0.125 in. Open-ended Pile Segment Model . . . . 1847.5.2 The 3 in. x 0.065 in. Open-ended Pile Segment Model . . . . 204

7.6 Pile S e t u p ..........................................................................................................204

8 NUMERICAL EXPERIMENTS ON A FULL-SCALE CLOSED- ENDED P I L E .................................................................................................. 2078.1 In troduction ...................................................................................................... 2078.2 Pile C haracteristics......................................................................................... 2078.3 Simulation of Pile D r iv in g .............................................................................2098.4 Consolidation and Pile Load T e s ts ................................................................2098.5 Pile S e tu p ......................................................................................................... 213

9 SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS . 2189.1 Summary ......................................................................................................... 2189.2 C onclusions...................................................................................................... 2199.3 Recommendations............................................................................................ 220

R E FE R E N C E S........................................................................................................ 222

A PPEN D IX A: PILE SETUP RELATED ST U D IE S............................234

A PPEN D IX B: NUMERICAL SIMULATION OF THE BEHAVIOR OF 3 in. x 0.065 in. O PEN -ENDED PILE SEGMENT MODEL ....................................................... 242

V I T A ........................................................................................................................... 253

VI


LIST OF TABLES

3.1 Soil properties of normally consolidated Sabine clay (after Bogard and Matlock, 1979, and Earth Technology Corporation, 1986)...................... 46

3.2 Material constants for Sabine clay (after Wathugala, 1990).................... 47

5.1 Dimensions of the pile segment models used in the field experiments. 86

A .l Pile setup related studies ............................................................................... 235

vii


LIST OF FIGURES2.1 Full scale field data on bearing capacity increase with time of friction

piles driven into c la y ...................................................................................... 14

2.2 Dissipation of pore water pressure during soil consolidation aroundthe Piezo-Lateral Stress cell in Empire clay (after Azzouz, 1986). . . 16

2.3 Decrease in bearing capacity with time after driving friction piles intoc l a y . .................................................................................................................. 17

2.4 Comparison of measured increase in bearing capacity of pile with theoretical increase in strength of soil adjacent to (a) Pile driven into Drammen clay, (b) Pile driven into San Francisco Bay mud (after Randolph et al., 1979)...................................................................................... 23

2.5 Results of static tension load tests conducted on a 3 in. x 0.125 in.open-ended pile segment model at Sabine Pass, Texas (after Earth Technology Corp., 1986).................................................................................. 25

3.1 Shape of the yield surface in the Ji-y/JiD space...................................... 30

3.2 (a) Shape of the yield surface in the triaxial plane; (b) Trace of theyield surface in the octahedral plane............................................................ 31

3.3 The yield surface, F, and the reference surface, R, in the triaxial plane. 33

3.4 Results of direct simple shear test on NC Boston Blue Clay, (a) Stresspath; (b) Shear stress-strain and pore water pressure-strain curves (after Ahmad, 1990)......................................................................................... 38

3.5 Location of the translation tensor, ay, in the J \-y /J m space (after Wathugala, 1990).............................................................................................. 40

3.6 Location of the translation tensor, ay; section A-A in the octahedral plane (after Wathugala, 1990)........................................................................ 43

3.7 (a) Comparison of predicted effective stress paths during undrainedtriaxial compression test using and models; (b) The corresponding predicted stress-strain relationships...................................................... 49

3.8 (a) Variation of predicted effective stress path during undrained triaxial compression test with Aq\ (b) The corresponding predicted stress- strain relationships........................................................................................... 50

3.9 (a) Variation of predicted effective stress path during simple shear testwith A q; (b) The corresponding predicted stress-strain relationships. . 52

viii


3.10 Comparison of predicted and measured effective radial stresses immediately after driving 1.72 in. diameter pile segment model into Sabine clay...................................................................................................... 53

3.11 Comparison of predicted and measured effective radial stresses immediately after driving 3 in. diameter pile segment model into Sabine clay ................................................................................................................... 54

3.12 (a) Variation of effective stress patb with A, during pile driving aspredicted by the model, using strain path method; (b) The corresponding predicted stress-strain relationships.......................................... 56

3.13 Predicted distribution of effective radial stress for different A, values immediately after pile driving, using the model................................. 57

3.14 Predicted distribution of pore water pressure for different Aq values immediately after pile driving, using the * u model................................ 58

4.1 Geometry of the simple pile........................................................................... 65

4.2 Strain field around simple pile for (a) Soil element located initially atra = R, (b) Soil element located initially at r„ = 2R ................................ 68

4.3 Distribution of strains immediately after driving at a depth z = R. . . 69

4.4 Geometry of the open-ended pile................................................................. 72

4.5 The effect of pile driving velocity, U, on the strain field of the surrounding soil (soil element is located at radial distance r = 1.5i2). . . 76

4.6 Strain field predicted by SPM due to an open-ended pile driving; (a)Soil element outside the pile located finally at radial distance r / =1.5f2, (b) Soil element inside the pile located finally at radial distanceTf = 15f............................................................................................................... 77

4.7 Strain paths, in the deviatoric space, of soil elements located at theimmediate vicinity outside and inside the wall during open-ended pile driving; (a) Ei vs E2] (b) E3 vs E2.............................................................. 79

4.8 The predicted effective stress path for a soil element at the pile-soilinterface path during pile driving................................................................. 81

4.9 Predicted distribution of effective stresses immediately after pile driving. 82

4.10 Predicted distribution of pore water pressure immediately after pile driving............................................................................................................... 83

5.1 The test procedure sequence for the field experiments........................... 87

ix


5.2 The 1.72 in. diameter pile segment model (after Earth TechnologyCorp., 1986)....................................................................................................... 89

5.3 The 3 in. diameter pile segment model (after Earth Technology Corp.,1986).................................................................................................................... 90

5.4 Variations of the measured pressures during soil consolidation aroundthe 1.72 in. diameter pile segment model (after Earth Technology Corp., 1986)....................................................................................................... 91

5.5 Results of the field static tension tests conducted on the 1.72 in. diameter pile segment model (after Earth Technology Corp., 1986). . . 93

5.6 Diagram for the use of UMAT in the finite element analysis................. 97

5.7 ABAQUS axisymmetric solid element CAX8RP used in the finite element analysis................................................................................................... 98

5.8 The finite element mesh used to discretize the soil domain around the1.72 in diameter pile segment model..............................................................100

5.9 Variation with radial distance of predicted effective stress paths duringpile driving for soil elements located 16 m below the ground surface. . 101

5.10 Predicted effective stress path for a soil element at the soil-pile interface during driving the 1.72 in. diameter pile segment model. . . 103

5.11 Predicted distribution of effective stresses at 16 m below the ground surface immediately after driving 1.72 in. diameter pile segment model. 104

5.12 Comparison of predicted and measured effective radial stresses duringsoil consolidation around the 1.72 in. diameter pile segment model. . 106

5.13 Predicted distribution of pore water pressure immediately after driving 1.72 in. diameter pile segment model.....................................................107

5.14 Sequence of the load tests conducted during soil consolidation aroundthe 1.72 in. diameter pile segment model.....................................................109

5.15 Comparison of predicted and measured normalized excess pore waterpressures during soil consolidation around the 1.72 in. diameter pile segment model....................................................................................................I l l

5.16 Predicted dissipation of pore water pressures during soil consolidationaround the 1.72 in. diameter pile segment model.......................................112

5.17 Predicted increase in effective radial stresses during the consolidationof the soil around the 1.72 in. diameter pile segment model....................113

x


5.18 Predicted distribution of effective stress at a depth of 16 m at the end of soil consolidation around the 1.72 in. diameter pile segment model....................................................................................................................114

5.19 Predicted effective stress path for a soil element at the pile-soil interface during driving and consolidation of the soil around the 1.72in. diameter pile segment model.................................................................... 116

5.20 Comparison of predicted and measured response during pile load test# 1 conducted on the 1.72 in. diameter pile segment model...........117

5.21 Comparison of predicted and measured response during pile load test# 2 conducted on the 1.72 in. diameter pile segment model...........119

5.22 Comparison of predicted and measured response during pile load test# 3 conducted on the 1.72 in. diameter pile segment model...................121





5.27 Comparison of predicted and measured response (shear transfer vs displacement) during the final cyclic pile load test conducted on the1.72 in. diameter pile segment model.......................................................... 132

5.28 Comparison of predicted and measured response (shear transfer vs time) during the final cyclic pile load test conducted on the 1.72 in. diameter pile segment model...........................................................................133

5.29 Comparison of predicted and measured response during the finalcyclic pile load test conducted on the 1.72 in. diameter pile segment model....................................................................................................................134

5.30 Comparison of predicted and measured increase in soil-pile shear transfer at failure (skin friction) for pile load tests performed at different time periods after driving.....................................................................136

6.1 Variation of measured pressures during soil consolidation around the3 in. diameter pile segment model (after Earth Technology Corp., 1986). 139

xi


6.2 Results of the static tension load tests conducted on the 3 in. diameterpile segment model...............................................................................................140

6.3 Sequence of the load tests conducted during soil consolidation on the3 in. diameter pile segment model................................................................... 141

6.4 The finite element mesh used for discretize the soil domain around the3 in. diameter pile segment model................................................................... 142

6.5 Predicted effective stress path for a soil element at the soil-pile interface during driving 3 in. diameter pile segment model................................144

6.6 Predicted distribution of effective stresses at 16 m below the ground surface immediately after driving the 3 in. diameter pile segment model. 145

6.7 Comparison of predicted and measured effective radial stresses after driving and during soil consolidation around the 3 in. diameter pile segment model...................................................................................................... 146

6.8 Predicted distribution of pore water pressure immediately after driving3 in. diameter pile segment model................................................................... 147

6.9 Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 3 in. diameter pile segment model...................................................................................................... 149

6.10 Predicted distribution of pore water pressures during consolidationof the soil around the 3 in. diameter pile segment model......................... 150

6.11 Predicted increase in the distribution of effective radial stresses during consolidation of the soil around the 3 in. diameter pile segment model....................................................................................................................151

6.12 Predicted distribution of effective stresses at a depth of 16 m at theend of soil consolidation around the 3 in. diameter pile segment model. 152

6.13 Predicted effective stress path for a soil element at the soil-pile interface during driving and soil consolidation around the 3 in. diameterpile segment model............................................................................................ 154

6.14 Comparison of predicted and measured response during pile load test# 1 conducted on the 3 in. diameter pile segment model.........................155



xii


6.17 Comparison of predicted and measured response during pile load test# 4 conducted on the 3 in. diameter pile segment model........................ 162

6.18 Comparison of predicted and measured response during the cyclicpile load test conducted on the 3 in. diameter pile segment model. . 164

6.19 Comparison of predicted and measured increase in soil-pile shear transfer (setup) during pile load tests performed at different timeperiods after driving......................................................................................... 168

7.1 Variation of measured pressures during soil consolidation around the 3 in. x 0.125 in. open-ended pile segment model driven at Sabine Pass, Texas (after Earth Technology Corp., 1986)................................... 171

7.2 Variation of measures pressured during soil consolidation around the 3 in. x 0.065 in. open-ended pile segment model driven at Sabine Pass, Texas (after Earth Technology Corp., 1986)................................... 172

7.3 Sequence of the load tests conducted on the 3 in. x 0.125 in. open- ended pile segment model..................................................................................173

7.4 Sequence of the load tests conducted on the 3 in. x 0.065 in. open- ended pile segment model..................................................................................174

7.5 Results of tension pile load tests conducted on the 3 in. x 0.125 in. open-ended pile segment model (after Earth Technology Corp., 1986). 175

7.6 Results of tension pile load tests conducted on the 3 in. x 0.065 in. open-ended pile segment model (after Earth Technology Corp., 1986). 176

7.7 The finite element mesh used to simulate the soil behavior around the3 in. x 0.125 in. and 3 in. x 0.065 in. open-ended pile segment models.177

7.8 Predicted effective stress path for a soil element at the soil-pile interface during driving and soil consolidation around the 3 in. x 0.125 in. open-ended pile segment model....................................................................... 180

7.9 Predicted distribution of effective stresses at 16 m below the groundsurface immediately after driving 3 in. x 0.125 in. open-ended pile segment model..................................................................................................... 181

7.10 Comparison of measured and predicted effective radial stresses during soil consolidation around the 3 in. x 0.125 in. open-ended pile segment model................................................................................................... 182

7.11 Variation with radial distance of predicted pore water pressure immediately after driving 3 in. x 0.125 in. open-ended pile segment model................................................................................................................... 183

xiii


7.12 Predicted dissipation of pore water pressures during soil consolidation around the 3 in. x 0.125 in. open-ended pile segment model................. 185

7.13 Predicted increase in effective radial stresses during soil consolidation around the 3 in. x 0.125 in. open-ended pile segment model................... 186

7.14 Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 3 in. x 0.125 in. open- ended pile segment model.................................................................................187

7.15 Predicted distribution of effective stresses at a depth of 16 m at the end of soil consolidation around the 3 in. x 0.125 in. open-endedpile segment model.............................................................................................189

7.16 Comparison of predicted and measured response for pile load test #1 conducted on the 3 in. x 0.125 in open-ended pile segment model. 190





7.21 Predicted response during pile load test # 5 performed on the 3 in.x 0.125 in. open-ended pile segment model............................................. 200

7.22 Comparison of predicted and measured response for the final cyclic pile load test conducted on the 3 in. x 0.125 in open-ended pile segment model....................................................................................................201

7.23 Predicted response for the final cyclic pile load test conducted on the3 in. x 0.125 in open-ended pile segment model........................................203

7.24 The finite element prediction of the increase in shear transfer for pile load tests conducted during soil consolidation around the 3 in. x 0.125 in open-ended pile segment model......................................................205

7.25 Comparison of the predicted and measured pile setup for the 3 in. x 0.125 in open-ended pile segment model......................................................206

8.1 The finite element mesh used to discretize the soil round the pile. . . 208

xiv


8.2 Predicted distribution of effective stresses immediately after driving0.5 m diameter full-scale pile............................................................................ 210

8.3 Predicted distribution of pore water pressure immediately after driving0.5 m diameter full-scale pile............................................................................ 211

8.4 Variation with time of predicted pore water pressure during soil consolidation around the 0.5 diameter full-scale pile.........................................212

8.5 Variation with time of predicted effective radial stress during soil consolidation around the 0.5 diameter full-scale pile...................................... 214

8.6 Variations with displacement of average skin friction along the shaft for different pile load tests conducted during soil consolidation aroundthe 0.5 m diameter full-scale pile..................................................................215

8.7 Predicted increase in average skin friction along the shaft of the pile during soil consolidation around the 0.5 m diameter full-scale pile. . . 216

B .l Comparison of measured and predicted effective radial stresses during soil consolidation around the 3 in. x 0.065 in. open-ended pile segment model..................................................................................................243

B.2 Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 3 in. x 0.065 in. open- ended pile segment model.............................................................................. 244

B.3 Predicted distribution of effective stresses at 16 m below ground surface immediately after driving the 3 in. x 0.0.065 in. open-ended pile segment model and subsequent consolidation.................................... 245

B.4 Variation of predicted pore water pressure with normalized radial distance immediately after driving the 3 in. x 0.065 in. open-ended pile segment model..................................................................................................246

B.5 Variation of predicted effective radial stress with normalized radial distance during soil consolidation around the 3 in. x 0.065 in. open- ended pile segment model.............................................................................. 247

B.6 Variations of predicted pore water pressure distribution with normalized radial distance during soil consolidation around the 3 in. x 0.065 in. open-ended pile segment model............................................................. 248

B.7 Comparison of the predicted and measured shear transfer vs displacement for pile load test # 1 conducted on the 3 in. x 0.065 in open- ended pile segment model.............................................................................. 249

xv


B.8 Comparison of the predicted and measured shear transfer vs displacement for pile load test # 2 conducted on the 3 in. x 0.065 in open- ended pile segment model................................................................................. 250

B.9 Comparison of the predicted and measured shear transfer vs displacement for pile load test # 3 conducted on the 3 in. x 0.065 in open- ended pile segment model................................................................................. 251

B.10 Comparison of the predicted and measured shear transfer vs displacement for pile load test # 4 conducted on the 3 in. x 0.065 in open-ended pile segment model.................................................................... 252

xvi


ABSTRACT

A general method for predicting pile setup, or the increase in the load carrying

capacity of friction piles driven into saturated clay, was developed. The solution is

based on the HiSS modeling approach, the strain path method, and the theory of

nonlinear porous media. The ability of this method was demonstrated by predicting

the field behavior of pile segment models during driving, subsequent consolidation,

and load testing. Numerical simulations were performed on pile segment models

because of the availability of field experimental tests necessary for verification and

evaluation schemes. The method was also applied to investigate the behavior of

full-scale friction piles driven into saturated clay.

An HiSS constitutive interface model named was developed in order to capture

soil behavior under severe shear deformation at the soil-pile interface. Verification

and evaluation of the model demonstrated the ability of the model to characterize

soil behavior at the soil-pile interface.

Simulation of pile driving was achieved by the strain path method. The simple pile

approach was used for full-displacement (closed-ended) piles, while the concept of

the ‘ideal’ open-ended pile was utilized for partial-displacement (open-ended) piles.

Numerical simulations of pile driving were performed and successfully predicted the

field behavior of the pile segment models during the driving.

Finite element simulation of soil consolidation around the pile was conducted.

During the consolidation phase, pile load tests were simulated at different time in

tervals. The purpose of simulating these tests was to predict the variation of pile

shaft capacity with time. These simulations were conducted using full-displacement

(closed-ended) as well as partial-displacement (open-ended) pile segment models.

xvii


The results of the finite element analysis were consistent with the measurements in

the field.

Investigating the behavior of the pile during its various life stages resulted in the

reasonable evaluation of pile shaft capacity. Finite element simulation of different pile

load tests during soil consolidation provided quantitative measures for the variation

in pile shaft capacity over time.

xviii


CHAPTER 1 INTRODUCTION

1.1 Background

Based on the nature of support provided by the surrounding soil, piles may be clas

sified as end-bearing piles and friction piles. While end-bearing piles transfer most

of their loads to an end-bearing stratum, friction piles resist a significant portion of

their loads via the skin friction developed along the surface of the pile. The behavior

of friction piles mainly depends on the interaction between the surrounding soil and

the pile shaft.

Over the past th irty years, investigators have studied the skin friction of piles

driven into cohesive soil, in order to estimate the pile carrying capacity. Different so

lutions have been proposed for this problem, most of which are empirical approaches

wherein skin friction is correlated to soil characteristics. The shaft friction can be

estimated using the following empirical methods: (a) a-m ethod (Peck, 1958; Wood

ward et al., 1961; Tomlinson, 1971; Flaate, 1972; McClleland, 1972), in which the

skin friction is correlated to the undrained shear strength of the undisturbed soil,

factor a accounting for the soil disturbance caused by pile installation; (b) /3-method

(Zeevaert, 1959; Eide et al., 1961; Chandler, 1968; Burland, 1973; Meyerhof, 1976;

Flaate and Seines, 1977; Parry and Swain, 1977), in which the skin friction is corre

lated to the initial effective overburden stress; (c) A-method (Vijayvergiya and Focht,

1972; Kraft et al., 1981), in which the skin friction is related to both undrained shear

strength and initial effective overburden stress; and (d) p-method (Azzouz et al.,

1990), in which the skin friction is correlated to the horizontal effective stress acting

on the pile shaft at the end of consolidation.

1


2

The bearing capacity (means load, carrying capacity) of friction piles driven into

clay may increase with time after driving. Sometimes called pile setup (or freeze), this

phenomenon has been observed by many researchers in different parts in the world

(Yang, 1956; Seed and Reese, 1955; Housel, 1958; McClelland, 1969; McCammon

and Golder, 1970; Flaate, 1972; Thorbum and Rigden, 1980; Konard and Roy, 1987;

McManis et al., 1988; Fellenius et al., 1989; Skov and Denver 1988). However, in a

few cases (Kraft et al., 1981; Cox and Kraft, 1979; Fellenius et al., 1989; Jardine and

Bond, 1989) the bearing capacity has decreased with time after pile driving.

Generally, current friction pile design is carried out based on empirical formu

las and depends to a large extent on the personal experience and judgment of the

engineer. Because of many uncertainties associated with pile foundation analysis

and design, full scale pile load tests are usually carried out at the site for important

projects. It is customary to do these tests 2 to 3 weeks after driving. The purpose

of a pile load test is mainly to assist the design engineer and to provide actual eval

uation of the pile response under loading. But these tests cannot substitute for the

engineering analysis of the pile behavior.

The empirical design methods do not provide a ‘general solution’ but rather a so

lution for a given site, pile characteristics, and driving and loading conditions. More

over, the existing design methods are conservative and may lead to over designed-

piles. Piles are expensive structural members, and pile projects are always costly.

For example, LaDODT (Louisiana Department of Transportation and Development)

spent about $7.8 million for precast concrete driven piles in Louisiana in 1993

(LaDODT, 1994). In the engineering practice, pile setup that develops after the

time of performing the load test is more often than not ignored. Therefore, a reli

able pile design based on a rational method that accounts for the time effect on the

shaft capacity of piles driven into saturated clay may reduce the cost and provide

the required performance for the pile.


However, few researchers (Randolph et al., 1979; Bogard and Matlock, 1990) have

investigated pile setup during their main studies of the behavior of driven piles. Pile

setup has been described generally in geotechnical engineering literature. Sometimes

it is mentioned incidentally through the discussion of the main points of interest. But

at present, there is no systematic research devoted solely to studying the variation

of the bearing capacity with time of friction piles driven into saturated clay.

1.2 O bjectives and Scope o f the Research

A reliable prediction of the shaft capacity of friction piles driven into clay is com

plicated because of the effects of pile installation on the mechanical behavior of the

soil. A rational and reliable technique for predicting the maximum shaft capacity

of friction piles driven into saturated clay, as well as the variation of this capacity

with time, is proposed. A consistent formulation throughout pile driving, subse

quent consolidation, and loading is used. Understanding of the mechanical behavior

of the soil-water system around the pile during driving, subsequent consolidation,

and loading is essential for achieving this goal. This method could provide reliable

and practical solutions to a wide range of soil conditions, pile characteristics, and

driving and loading conditions. Moreover, incorporating pile setup predicted by this

method in the design procedures will lead to an economical and safe design and will

also decrease the dependence on the personal judgment of the engineer.

The objectives of this research are:

1. to develop a rational method in order to predict the mechanics involved in pile

setup, or the increase in the shaft capacity with time of friction piles driven

into saturated clay.

2. to develop an elasto-plastic constitutive model, based on the HiSS modeling

approach, to characterize the behavior of clay at the soil-pile interface.


3. to determine the strain field of the soil around the pile during driving (based

on the strain path method) for partial-displacement (open-ended) piles.

4. to implement the proposed model into the finite element program ABAQUS,

in which the coupled theory of nonlinear porous media is used for the analysis.

5. to verify and evaluate the proposed method using the finite element simulation

of the field behavior of a well-documented case history.

6. to perform finite element simulations of pile load tests in order to estimate pile

setup and compare results with field measurements. The simulations include

closed-ended as well as open-ended piles.

7. to perform numerical experiments in order to study the setup behavior of a

full-scale pile.

1.3 Organization o f the Manuscript

The manuscript is organized into nine chapters. A brief summary of their contents

is presented in this section.

In Chapter One, the significance of the current study is discussed, and the objec

tives and scope of this research sure presented. A comprehensive review of pile setup

literature is given in Chapter Two. The factors affecting soil-pile interaction during

pile installation, consolidation, and loading are analyzed. The pile setup phenomenon

is thoroughly reviewed and interpreted. Furthermore, pile setup related studies are

compiled and presented in tabular form. The development of the ^‘-series of the

Hierarchical Single Surface modeling approach is covered in Chapter Three. The

proposed HiSS-cS^ model is presented, and the predictive capability of the proposed

model is discussed. Implementation of the proposed model into a computer code as

well as into the finite element program ABAQUS is presented. In Chapter Four, the


5

simulation, of pile driving utilizing the strain path method is reviewed. The strain

field in the soil around the pile due to driving full-displacement (closed-ended) piles

as well as partial-displacement (open-ended) piles is determined. The verification of

the numerical procedure utilizing the Sabine Pass case study is presented in Chapter

Five. Finite element simulations of the consolidation and pile load tests during the

consolidation are conducted. Pile setup is determined. Discussion of the verification

scheme and its results are also covered. In Chapter Six, numerical simulations of pile

driving, consolidation, and pile load tests are performed on another closed-ended

pile segment model. Comparisons of the predicted and measured pile response are

presented. The numerical simulations of the behavior of open-ended pile segment

models are covered in Chapter Seven. Also, numerical results are compaxed with

field measurements during pile driving, consolidation, and load testing. In Chapter

Eight, numerical experiments are conducted on a full-scale pile. The behavior during

pile driving, consolidation and load testing is analyzed. Finally, in Chapter Nine,

the summary, conclusions, and recommendations are presented.


CHAPTER 2 BACKGROUND

2.1 Introduction

Piles axe relatively long and generally slender structural foundation members that

transmit superstructure loads to deep soil layers. In geotechnical engineering, piles

usually serve as foundations when soil conditions are not suitable for the use of

shallow foundations. Moreover, piles have other applications in deep excavations

and in slope stability. As presented in the literature, piles axe classified according to:

• the nature of load support (friction and end-bearing piles),

• the displacement properties (full-displacement, partial-displacement, and non

displacement piles), and

• the composition of piles (timber, concrete, steel, and composite piles).

The behavior of the pile depends on many different factors, including pile charac

teristics, soil conditions and properties, installation method, and loading conditions,

among others. The performance of piles affects the serviceability of the structure

they support. In turn, the different life stages of the pile affect its performance. It is

important, therefore, to investigate and understand the behavior of the pile during

its various life stages.

In the current study, the behavior of friction piles driven into saturated clay is

investigated. Friction piles resist the load through the friction develops along the

shaft. In order to study thoroughly the behavior of friction piles, the effects of pile

installation on the surrounding soil should be analyzed. The stages that a driven

pile experiences during its life are:

6


7

• pile driving,

• changes in the stress field as well as in pore water pressures in the surrounding

soil,

• equalization of the excess pore pressures and the stress field in the soil around

the pile, and

• pile loading, which includes load tests and/or construction of the proposed

structure.

In this chapter, a review of the installation effects of a driven pile on the clay-

water system is presented. Case studies, which investigate time dependent changes

in shaft capacity for friction piles driven into clay, are compiled and analyzed. The

results of the literature review are su m m arized in Table A .l in Appendix A.

2.2 Effects o f P ile Driving on the Clay—W ater System

Pile driving considerably affects the surrounding soil. Understanding and analyzing

the effects of pile driving on the surrounding soil is thus necessary to improve the

estimation of the bearing capacity of the pile. Many researchers have studied these

effects and their variations, experimentally as well as theoretically, through the var

ious life stages of the pile. According to de Mello (1969), the effects of pile driving

on the clay-water system are:

1. remolding and disturbance (partial structural alteration) of the soil surrounding

the pile,

2. changes in the stress state of the soil in the vicinity of the pile,

3. dissipation of the excess pore water pressure generated in the soil surrounding

the pile, and


8

4. regain of the soil strength over time.

These effects are discussed in detail below.

2.2.1 D isturbance Effects on Soil Properties

The amount of soil disturbance around a driven pile, and its effects on soil prop

erties and pile behavior, have been widely investigated (Housel and Burkey, 1948;

C u m m in gs et al., 1950; Seed and Reese, 1955; Orrje and Broms, 1967; Flaate, 1972;

Fellenius and Samson, 1976; Bozozuk et al., 1978). Results indicated that an ap

preciable decrease in the undrained shear strength of the surrounding soil occurred

because of remolding due to driving. In some cases, the reduction occurred in a zone

around the pile that extended up to twice the pile diameter away from the surface

of the pile. Terzaghi (1941) attributed the decrease in soil strength upon remolding

to the destruction of particle contacts which allows more water to fill in old contact

points.

Changes in the water content of the soil after pile driving have been observed by

Cu m m in gs et al. (1950), Seed and Reese (1955), and Flaate (1968). Seed and Reese

(1955), performed a study on the behavior of friction piles driven into San Francisco

Bay mud. They reported a progressive reduction of the water content of the soil near

the pile after driving. A detailed study, conducted by Flaate (1968) at the Nitsund

bridge site in Norway, showed a significant decrease in soil water content in a zone

that extended up to 15 cm from the pile surface.

2.2.2 Changes in Pore W ater Pressures

When a pile is driven into soil, it displaces a volume of soil equal to its own volume.

Ground heave is associated with pile driving at small penetrations up to five times

pile diameter (Cooke and Price, 1973). At greater depths, the soil movement occurs

primarily in the radial direction. Pile driving into saturated clay is usually assumed


9

to occur under undrained conditions because the process is performed in a relatively

short time.

The change in pore water pressure due to pile driving is attributed to the change

in the stress state of the surrounding soil. According to the effective stress principle,

the pore water pressure is equal to the difference between the total and effective

stresses. During pile driving, an increase in the total mean stress occurs as the pile

forces the soil out of its path. Moreover, severe shearing and remolding of the soil

in the vicinity of the pile may cause either an increase or a decrease in the effective

stresses, depending on whether the soil tends to dilate or contract upon shearing.

In normally consolidated and slightly overconsolidated clays, a decrease in effective

stresses occurs as the soil tends to contract upon undrained shearing. The net result

is an increase in the pore water pressure.

On the other hand, for a heavily overconsolidated clay, negative pore water pres

sures are generated due to the increase in the effective stresses, as soil tends to dilate

during undrained shearing. At the same time, positive pore water pressures develop

due to the increase in total mean stress as a result of driving. When the negative

part of the pore water pressures dominates, it is expected that a net decrease in pore

water pressures due to pile driving will be induced.

High excess pore water pressure buildup in clay after pile driving was observed

by Bjerrum et al. (1958), Lo and Stermac (1965), Lambe and Horn (1965), Koizumi

and Ito (1967), Orrje and Broms (1967), Flaate (1968), D’Appolonia and Lambe

(1971), Clark and Meyerhof (1972), Bozozuk et al. (1978), Sutton and Rigden (1979),

Konard and Roy (1987), Chung (1988), Bogard and Matlock (1990), Leung et al.

(1991) and many other investigators. Results indicated that the pressure decreased

with the increase of the radial distance from the pile surface. The radial effect

extended up to ten times the pile diameter. The magnitude of excess pore water


10

pressure at a certain depth approached the initial total overburden pressure, and in

many cases even exceeded this amount.

Many researchers have proposed solutions to predict the induced pore water pres

sures due to driving. Theoretical solutions were proposed by Nishida (1964) based

on the elastoplasticity analysis, and by Ladanyi (1963), using the expansion of a

cylindrical cavity to estimate the pore water pressure distribution in the soil around

the pile. Lo and Stermac (1965) presented a theoretical solution that estimates the

maximum pore water pressure induced by pile driving. This method is applicable to

normally consolidated and slightly overconsolidated clays. The solution is based on

the assumption that the pore water pressure increases as a result of the increase in

the mean total stress and shearing of the soil around the pile. The results obtained by

this method were consistent with the field measurements. D’Appolonia and Lambe

(1971) derived another form of the Lo and Stermac solution. Randolph et al. (1979)

introduced an expression to estimate the excess pore water pressure develops close

to the pile. The formula relates the excess pore water pressure to the change in

the mean effective stress due to remolding and shearing and to the undrained shear

strength of the soil.

However, very little information is available about the generation of negative

excess pore water pressure due to driving. The reason might be the postulate that

negative pore water pressure is induced in heavily overconsolidated clay. This clay is

stiff and strong and may be suitable for other types of foundations. Field experiments

were conducted by Jardine and Bond (1989) on instrumented closed-ended steel

piles installed in heavily overconsolidated London clay. Measurements of pore water

pressure along the pile shaft during pile installation showed generally negative values.

Jardine and Bond (1989) and Bond and Jardine (1991) stated that the soil dilated

during installation and that significant increases in the effective radial stress and

negative pore water pressure were observed.


11

2.2.3 Equalization o f Pore W ater Pressures and Stresses

As soon, as changes in pore water pressures occur, the equalization phase begins.

The process continues as pore water pressures change over time, as a result of pore

water migration from high excess pressure to low excess pressure zones. This is

accompanied by changes in the stress field of the soil around the pile.

In the case of normally consolidated and slightly overconsolidated clays, the ex

cess pore water pressure dissipates mainly by radial outward flow resulting in soil

consolidation. An increase in the effective stress field in the soil around the pile

occurs during the consolidation stage.

Soderberg (1962) presented a solution for the rate of dissipation of excess pore

water pressure around a driven pile. Dissipation was assumed to occur by radial

flow of pore water away from the pile. Vertical dissipation that might have occurred

around the tip and the top of the pile was neglected.

At the Nitsund bridge site study, a remarkable increase in the shear strength of

the soil at the end of consolidation (as compared to its original strength) was reported

by Flaate (1972). Orrje and Broms (1967) observed an increase in undrained shear

strength of soil around a pile nine months after driving. Seed and Reese (1955)

measured an increase in the shear strength of the soil from an initial value of 0.24

kg/cm 2 before driving to a 0.36 kg/cm2 thirty days after driving. The increase was

associated with a decrease in the water content from 48.1% to 41.1%.

The increase in the shear strength is attributed to the increase in the effective

stresses in the soil around the pile. Soderberg (1962) attributed the increase, over

time, in the bearing capacity of friction piles driven into clays and silts, to the

subsequent consolidation. Another phenomenon, known as thixotropy, contributes

to the increase in the strength of the sensitive clays. The undrained strength of the

clay is regained as a large percentage of the destroyed structural bonds due to soil

remolding is reestablished.


12

In heavily overconsolidated clays, an inward flow of the pore water occurs due to

the generation of negative pore water pressures. The effective stresses decrease as

the process continues, and, with time, swelling takes place in the soil surrounding

the pile. As a result, soil strength decreases in addition to the strength reduction

from soil remolding. Therefore, over time, a decrease in the shaft capacity of friction

piles driven into heavily overconsolidated clay is expected. Jardine and Bond (1989)

reported that both swelling and consolidation were involved in the equalization pro

cess in heavily overconsolidated London clay. By the end of the equalization process,

a slight decrease in the bearing capacity was observed.

Due to lack of knowledge and limited field studies regarding the behavior of

overconsolidated clays, the preceding discussion is hypothetical.

2.3 P ile Setup

An extensive literature review was conducted in this study in order to collect the

available field measurements regarding pile setup. Cases of normally and slightly

overconsolidated clays were emphasized and are summarized in Table A .l in Ap

pendix A. The objectives of the literature review were to present the state-of-the-art

knowledge about pile setup and to locate a well-documented case history. This case

history was necessary in order to verify and evaluate the theoretical and numerical

approach proposed in this study.

Based on field observations, pile setup was reported by Yang (1956), Seed and

Reese (1955), Housel (1958), Eide et al. (1961), McClelland (1969), McCammon and

Golder (1970), Flaate (1972), Thorbum and Rigden (1980), Samson and Authier

(1986), Earth Technology Corp. (1986), Konard and Roy (1987), McManis et al.

(1988), Chung (1988), Skov and Denver (1988), Fellenius et al. (1989), Coop and

Wroth (1989) and Bogard and Matlock (1990). Most of these studies demonstrated


13

a time dependent increase in the bearing capacity of friction piles driven into clay

(Figure 2.1).

The increase in pile capacity with time was reported in different regions, includ

ing Louisiana. Some of the field investigations that were conducted in Louisiana are

discussed below. Blessey and Lee (1980) observed an approximately 400 to 500 per

cent increase in shaft capacity a few weeks after pile driving in southeast Louisiana.

McManis et al. (1988) presented the results of a field investigation concerning pile

bearing capacity determination. In the investigation, square and cylindrical full scale

concrete piles were driven near the 1-310 Luling Bridge near New Orleans, Louisiana.

These piles were subjected to static load tests and dynamic field measurements using

the Pile Driving Analyzer (Goble et al., 1980). Field measurements were obtained

to determine the variation of the bearing capacity of driven piles over time. The end

of driving bearing capacity was estimated based on the dynamic analysis performed

using the Pile Driving Analyzer. Results indicated that, in the first five weeks after

driving, the bearing capacity increased by a factor ranging from 4.4 to 11.5. The

variation of these capacities with time is included in Figure 2.1.

Konard and Roy (1987) performed a full-scale study on two 22 cm diameter

closed-ended steel pipe piles at a test site in St. Alban, Canada, to investigate pile

setup. Results showed a significant time dependent increase in the bearing capacity

of short friction piles jacked into overconsolidated (OCR=2.20) soft sensitive marine

clay. The increase reached twelve times the initial shaft capacity. The excess pore

water pressure generated during driving was 1.6 times the overburden pressure, and

it dissipated in about 600 hours.

As shown in Figure 2.1, the bearing capacity of friction piles driven into clay

may increase as much as twelve times the capacity at the end of driving. The

rate of increase in the pile setup is variable. In some cases, most of the setup was

developed rapidly during the short time period (about thirty days) after driving, and


Reproduced

with perm

ission of the

copyright ow

ner. Further

reproduction prohibited

without

permission.

o(0Q.80)

E3E'8

ca>Ea>CL

Seed and R eese, 1955

Yang, 1956

Yang, 1956

Housel, 1958

McClleland, 1969

Flaate, 1972

Flaate, 1972

Thorburn and Rigden, 1980

Konard and Roy, 1987

McManis, 1988

McManls, 1988

Skov and Denver, 1988

Fellenius et al., 1989

1.0E-1 1.0E+0 1.0E+1 1.0E+2

Time after driving, days

1.0E+3

Figure 2.1: Full scale field data on bearing capacity increase with time o f friction piles driven into clay.

15

then leveled off thereafter. Moreover, the pile setup curves are approximately the

inverse of the pore water pressure dissipation curves. Figure 2.2 shows a pore water

pressure dissipation curve. This indicates that the setup may be directly related to

the dissipation of the excess pore water pressures. Vesic (1977) presented field data

gathered from five cases in which the increase in the bearing capacity was reported

as 4 to 5 times the capacity at the end of driving. These case are included in Figure

2 . 1.

The results of a thorough review of pile setup related case studies are summarized

in Table A.I. Emphases were given to determine: (a) the soil conditions, (b) the

generation of the excess pore water pressure due to pile driving, and (c) the increase

in the bearing capacity with time after driving. Most of the soils shown in Table A .l

are normally consolidated, though some axe slightly over consolidated clays. Mea

surements of pore water pressure during pile driving, consolidation, and load testing

were conducted in some of these studies. Case studies in which measurements were

carried out showed an increase in pore water pressures due to pile driving. In most

of these studies, investigators reported an increase in pile bearing capacity with time

after driving (or pile setup). However, the decrease in the bearing capacity with

time after driving was also observed by Kraft et al. (1981), Fellenius et al. (1989),

and Bond and Jardine (1991). The measured decrease in bearing capacity with time

after driving is depicted in Figure 2.3. In these cases, no pore water pressure mea

surements were taken. Bond and Jardine (1991) performed field measurements on

friction piles driven into overconsolidated London clay. They showed that the effec

tive radial stress decreased during the pore water pressure equalization process, and

the long term pile capacities slightly decreased after jacking (i.e. negative pile setup

was observed).

The discussion presented in this Section indicates that the change in pile capacity

with time is significant. Within the presented case studies, the change was observed


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

1.2

3(A</>s>CL 0.8

ro£oaT>0)= 0.4(0§oz

0.0 i m0 1 10 100 1000 10000

Time factor, T=ct/r02

Figure 2.2: Dissipation of pore water pressure during soil consolidation around the Piezo-Lateral Stress cell in Empire clay (after Azzouz, 1986).

Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

600

Kraft et al. 1981, Empire site

Kraft et al. 1981, Empire site

Fellenius et al., 1989</>Q.V

500.S'on>Q.so>c•c(0a>

400a>K

300101 100 1000

Time after driving, Days

Figure 2.3: Decrease in bearing capacity with time after driving friction piles into clay.

18

mainly as an increase in the bearing capacity of the pile after driving (pile setup).

However, a decrease in bearing capacity after pile driving was also reported, although

it is not as common as pile setup. Therefore, a reliable method for predicting pile

setup would be very useful.

2.4 Soil Sensitivity

Clays exhibit reduction in strength upon remolding. The ratio of the undisturbed and

remolded strength of the clay is known as sensitivity. Different phenomena contribute

to the development of soil sensitivity. Among these are cementation, thixotropic

hardening, metastable structure and leaching, ion exchange, and change in mono

valent/divalent cation ratio. Mitchell (1993) defined thixotropy as am isothermal,

reversible, time dependent process occurring under conditions of constant composi

tion and volume, whereby a material stiffens while at rest and softens and liquefies

upon remolding. Mitchell (1960) also summarized previous studies of thixotropy

in suspensions and soils. He discussed the complex nature of and the important

factors involved in this phenomenon. These factors control the microstructural be

havior of the clay particles, such as pore fluid composition, particle size and shape,

and interaction between interparticle forces (attractions and repulsions). Based on

experimental observations, Mitchell (1960) proposed a hypothesis as a possible ex

planation of thixotropy effects in soils. But a comprehensive theory for the basic

cause of thixotropy is not available. Thixotropic regain is one of the factors that

cause soil sensitivity. This phenomenon occurs under controlled conditions such as

constant water content. Pore water pressure starts to dissipate immediately after

pile driving (actually it does so as soon as excess pore pressure is generated), leading

to a decrease in water content with time. Based on previous experiments, Skempton

and Northey (1952) showed that thixotropic regain decreases in accordance with the


decrease in water content below the liquid limit. In addition, they presented the re

sults of studies conducted on different clays (London, Shellhaven, and Detroit clay),

where 100% of the strength of the clay was regained after one year.

Many case studies of pile driving in insensitive soils showed an appreciable amount

of pile setup. For example, Seed and Reese (1955) reported that, th irty days after

pile driving, a setup of 5.4 times that at the end of driving occurred. The increase

was attributed to the increase in the shear strength of the surrounding soil. This

case of insensitive soil is compared (in this study) with the Skempton and Northey

(1952) study about sensitivity. It is concluded that the 100% strength regain, after

one yeax, due to thixotropy, is relatively small when compared to the increase of

540%, after one month, due to consolidation.

Based on these observations, it was concluded that the effect of thixotropy is small

(if there is an effect) compared to that due to consolidation. Moreover, the lack of

sufficient knowledge about thixotropy in soils was one of the reasons for not incor

porating this phenomenon in pile setup. Therefore, thixotropy was not taken into

account during this study, and pile setup was investigated based on consolidation.

2.5 M ethods of Predicting Pile Shaft Capacity

For many years, researchers have attem pted to explain, understand, and predict pile

shaft capacity, using various empirical, semi-empirical, theoretical, and numerical

techniques. However, there is still no rigorous theory that accounts for the condi

tions involved in the pile problem and that yields a ‘general solution’. Solutions for

specific conditions have been achieved, and discrepancies have appeared in field mea

surements and observations. Meanwhile, the understanding of soil-pile interaction

is improving.


20

Early interpretations (Peck, 1958; Woodward et al., 1961; Tomlinson, 1971;

Flaate, 1972; McCUeland, 1972) of the shaft-soil behavior were based on the to

tal stress approach, which correlates the strength of the soil-pile bond with the

undrained shear strength of the undisturbed soil.

Then the effective stress approach was introduced (Zeevaert, 1959; Eide et al.,

1961; Chandler, 1968; Burland, 1973; Meyerhof, 1976; Flaate and Seines, 1977; Parry

and Swain, 1977). It postulates that soil-pile behavior is governed by the effective

stress state, and that shaft capacity should be interpreted on this basis. Generally,

these studies were based on one-stage calculations and the life of the pile was not

considered.

Effective stress theories with treatment of all life stages of the pile then started

to appear. General effective stress theories (Esrig et al., 1977) for predicting pile

behavior during the different life stages as well as the bearing capacity for driven piles

were introduced. Kirby and Wroth (1977) utilized the critical state soil mechanics

concept to study pile behavior. Numerical and analytical methods, using cavity

expansion methods and the constitutive models to describe the soil response, were

developed to study pile behavior (Cater et al., 1979; Randolph et al., 1979; Randolph

and Wroth, 1979; Wroth el al., 1979). Instrumented piles, pile models, and pile

segment models were fabricated for experimental, theoretical, and numerical studies

(Banerjee et al., 1982; Roy et al., 1981; Azzouz and Lutz, 1986; Coop and Wroth,

1989; Bogard and Matlock, 1990; Bond and Jardine, 1991; Wathugala et al., 1993).

The results of these studies demonstrated the importance of the effective stress state

in the soil around the pile in evaluating pile shaft-soil behavior.

By considering the life stages of the pile, these studies treated pile behavior on the

basis of new considerations. Namely, proposing the constitutive laws to describe soil

behavior with the assistance of advanced laboratory testing programs; developing

ways to simulate pile driving, with the help of advanced computing power; and using


21

the finite element method and different numerical techniques to predict pile response

during consolidation and loading.

2.6 Pile—Setup Related Studies

Seed and Reese (1955), Eide et al. (1961), Flaate (1972), Banerjee et al. (1982),

Bogard and Matlock (1990), and many other investigators, as presented in Table A .l,

discussed the increase in pile hearing capacity with time after driving. During these

experimental studies, the changes in soil properties and/or the changes in stresses

and pore water pressures during the different life stages of the pile, were monitored.

Few theoretical and numerical studies were carried out to investigate these effects

(Randolph et al., 1979; Banerjee et al., 1982; Wathugala, 1990).

Randolph et al. (1979) performed a numerical analysis of the behavior of piles

driven into clay. Changes in stresses and pore water pressures during pile driving

and during the subsequent consolidation of the soil around the pile were investigated.

Pile installation was modeled as the undrained expansion of a cylindrical cavity.

The excess pore water pressures generated due to cavity expansion were assumed

to dissipate by outward radial flow of the pore water. CAMFE (CAMbridge one

dimensional Finite Element program) was used in the study. The finite element

method was utilized to study the changes in the stresses and pore water pressures

due to cavity expansion and during consolidation, with the soil modeled as modified

Cam-clay. This approach was used to estimate the changes in strength and water

content of soil adjacent to a driven pile. It was shown that the rate of increase of

the bearing capacity of a driven pile may be estimated from the predicted rate of

increase in shear strength of soil around the pile. Results of the pile setup predicted

by Randolph et al. (1979) are shown in Figure 2.4


22

The analysis was based on the assumption of plane strain and axial symmetry,

which reduced the problem to that of a one-dimensional category. Therefore, the

shear stress acting in the vertical-radial plane, in addition to the vertical movement

due to pile driving were not considered. The method proposed by Randolph et al.

(1979) predicts the pile capacity from the changes in the shear strength and water

content of the soil. Simulations of pile load tests and prediction of pile capacity were

not attem pted by Randolph et al. (1979).

Banerjee et al. (1982) conducted experimental, theoretical, and numerical studies

to investigate the behavior of axially loaded piles driven into saturated clay. The

stresses generated as well as the buildup and decay of the pore water pressures due

to driving an instrumented model pile into a normally consolidated kaolin bed, were

monitored continuously during and after pile driving. Theoretical and numerical

studies were performed to analyze the problem and to interpret the experimental

data. Pile driving was simulated as the penetration of a rigid, smooth, cylindrical

object within a soil mass, which was modeled as modified Cam-clay. This formu

lation Weis based on an axisymmetric updated Lagrangian formulation of the finite

element method. Another finite element code was developed to simulate the con

solidation process with consideration of small deformations. Finite element method

predictions of the pore water pressures at the soil-pile interface during pile driving

and subsequent consolidation were higher than those observed in the experiments.

A theoretical solution was also presented, based on the plane strain expansion of the

cylindrical cavity. The cavity expansion solution showed agreement with the exper

imental data at the soil-pile interface but disagreement with it at the boundaries.

This study provided small scale data used to calibrate numerical solutions to the

pile problem. During this study, no attempts were made to predict field data or to

simulate pile load tests.


23

125

100

83

O

3

O

Eide et al. (1961); field measurements

Randolph et al. (1979); finite element prediction

0 200 400 600 800


(a)

125

100

875 -

3

O

3

O

Seed and R eese (1955); field measurements

Randolph et al. (1979); finite element prediction

0 10 20 30


(b)

Figur 2.4: Comparison o f measured increase in bearing capacity of pile with theoretical increase in strength o f soil adjacent to (a) Pile driven into Drammen caly, (b) Pile driven into San Francisco Bay mud (after Randolph et al., 1979).


24

Through a joint study between the University of Arizona, Tucson, and the Earth

Technology Corporation (1986), field experiments were conducted at Sabine Pass,

Texas, using two instrumented pile segment models 1.72 in. and 3 in. (4.37 cm and

7.62 cm) in diameter with changeable cutting shoes. During installation, consoli

dation, and static and cyclic load tests, measurements of total radial stresses, pore

water pressures, shear transfer, and pile-soil displacement were carried out. Shown

in Figure 2.5 are the results of static tension tests performed on a 3 in. x 0.125 in.

(7.62 cm x 0.318 cm) pile segment model during consolidation. Wathugala (1990)

performed a theoretical and numerical study to investigate the behavior of piles

driven into saturated clay. A general procedure, based on finite element dynamic

analysis of nonlinear porous media, was developed to simulate and predict the re

sponse of the pile at its various life stages (initial conditions, driving, consolidation,

and load tests). The Hierarchical Single Surface modeling approach (Wathugala,

1990) was used to model clay behavior. The procedure successfully predicted the

fined cyclic load tests at the end of consolidation. Predictions of shear transfer versus

time and displacement were consistent with the field experiments. But this method

under predicted the pore water pressures immediately after driving and therefore

overestimated the effective stress state.


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

(0CL

& 20 CO C

2t_(00)-Cw U = 9 5 %

U = 86%

U = 57%

U = 36%

U = 16%

0.0E+0 1.0E-3 2.0E-3 3.0E-3 4.0E-3

Displacement, m

Figure 2.5: Results o f static tension load tests conducted on a 3 in. X 0.125 in. open-ended pile segment model at Sabine Pass, Texas ( after Earth Technology Corp., 1986).

to

CHAPTER 3 HIERARCHICAL SINGLE SURFACE

MODELING APPROACH

3.1 Introduction

Understanding the mechanical stress-deformation response of geologic materials to

external excitations is essential for obtaining reliable solutions to different geotechni-

cal engineering problems. Researchers have developed various mathematical theories

in order to understand and interpret the mechanical behavior of these materials.

Constitutive models, or laws, are the general terms used to describe these theo

ries. The most com m on constitutive models for geologic materials are based on the

mathematical theories of elasticity and plasticity.

The Hierarchical Single Surface modeling approach (HiSS), as proposed by Desai

and co-workers in 1980, 1986, and 1993, is used for geologic materials. This modeling

approach captures the behavior of solids as well as discontinuities such as interfaces

and joints. The behavior of a wide range of geologic materials such as soils, rocks,

and concrete has been predicted and verified with respect to both laboratory and field

experiments. A major characteristic of the HiSS modeling approach is that it allows

for the progressive development of models of higher grades corresponding to different

levels of complexity. These models remain under the general HiSS theory. Wathugala

(1990) developed the ^‘-series of the HiSS constitutive models to characterize the

behavior of cohesive soils. An interface model named based on the nonassociative

anisotropic 8 J model, is proposed in this study. This model, in addition to the general

description of the 5*-series of HiSS constitutive models, is presented in this chapter.

26


3.2 The (5*-Series o f HiSS Constitutive M odels

27

This series comprises a set of the HiSS models developed to study the behavior

of cohesive soil under different soil and loading conditions such as: normally and

overconsolidated soils; drained and undrained conditions; monotonic, cyclic, and

complex loading; and induced anisotropy.

The following is a general description of the HiSS ^‘-series models:

1. Sq model: the basic model with isotropic hardening and associative flow rale.

2. S* model: isotropic hardening and non associative flow rule.

3. 8 2 model: anisotropic hardening and non associative flow rale.

4. 5*+r model: strain softening, damage-based.

5 - s:+v model: viscoplastic, it is introduced using Perzyna’s theory.

3.2 .1 Basics o f the Increm ental Theory of Plasticity

The incremental theory of plasticity is based on the fundamental assumptions of

yield surface and of flow and hardening rules. These aspects describe the constitutive

behavior of the material as a result of external excitations.

The general incremental elastoplastic stress-strain relationship for any loading

may be given by

d<rH = °ijkid£ki (3.1)

or

dsij = D*jkldtrki (3.2)


28

the superscript * can be V L, R L , or UL, depending on virgin loading, reloading,

or unloading, respectively. Ciju and Diju are the elastoplastic constitutive stiffness

and compliance tensors, respectively. The general forms of these tensors are given in

Wathugala (1990) as

H* + n*C ;.tunl.

and

/'-»• _ /nre v i j m n 1Lm n 1,'opKjopkl f n< 'ijU - ~ r r , , o Q I*1 - 1 !

DiiU = D‘iiU + ^ (3.4)

where H* is the plastic modulus and Cfjki and Dfjki Me the elastic constitutive

stiffness and compliance tensors, respectively. The task of a constitutive model is

to define n^-, n - , and H* for all loading situations. The tensors nf- and to- are

defined as unit normal tensors to surfaces R and Q. In the case of the virgin loading,

H*(= H VL) is derived using the consistency condition, while for other cases, H*

(H RL and H UL) is obtained using the interpolation functions.

3.2.2 Y ield S urface , F

The yield surface, F, is defined in terms of the invariants of the stress tensor, crtJ-.

Desai and Wathugala (1987) proposed the nondimensional form of the yield function

as follows

F = ( 7 I ) - F"F- = 0 (3-5>

where J 2D is the second invariant of the deviatoric stress tensor, Pa is the atmospheric

pressure, and F& is the basic function given by:

A = - a . . ( £ ) + T ( A ) (3.6)


29

where J\ is the first invariant of the stress tensor and atp, is the hardening function.

The basic function, Ff,, describes the shape of F in the J \-y /J 2D space. The material

parameter n defines the phase change of the material from contraction to dilation,

while 7 is a material parameter related directly to the ultimate state and depends

on the angle of internal friction, <f>. The shape function, F5, describes the shape of

F in the octahedral plane and is given by:

F. = ( l - f 3 S r)m (3.7)

where /3 is the ultimate parameter that controls the shape of the yield surface in the

octahedral plane, m is a material parameter with a value equal to —0.5 for geologic

material (Wathugala, 1990), and Sr is the stress ratio given by:

_ yn J3p ( .

where J^d is the third invariant of the deviatoric stress tensor. The shape of the yield

surface, in addition to the phase change line and the ultimate line in the J \-y /J 2 D

space is shown in Figure 3.1. Moreover, Figure 3.2 depicts the shape of the yield

surface in the triaxial plane and the trace of this surface in the octahedral plane.

The HiSS ^’-series models treat virgin loading {VL), unloading (UL), and reload

ing (R L ) differently. When the material yields, the stress point lies on the yield

surface, F. If the stress increment, directs outward from the yield surface, then

virgin loading occurs. If the stress increment directs inward, then unloading occurs.

Neutral loading occurs when the stress increment is tangential to the yield surface,

F. In other words:

nf- d&ij > 0 virgin loading"tj “ '- 'u

lfj d<rHnf- daij < 0 unloading


009

30

o

CN

OW

O

C L

in*o

oooCN

O O

ed* ‘ azr ^


Figu

re

3.1:

Sha

pe

of the

yi

eld

surfa

ce

in the

J,

- V JID

sp

ace.

31

2000

Yield surface

HC line

0 1000 2000

V T a2= V T a3

(a)

(b)

Figure 3.2: (a) Shape o f the yield surface in the triaxial plane; (b) Trace of the yield surface in the octahedral plane.


32

nfj d<Tij = 0 neutral loading (3.9)

where nf- is the unit normal to the yield surface, F, and is given by

BFF _ n-- =... _ a<r* (3.10)3 [ BF BF I 1/ 2

L8<Tm»W hen the stress point lies inside the yield surface, F, it is important to distinguish

unloading from reloading. Therefore, the convex reference surface, R, which passes

through the current stress point in the stress space, is introduced and defined as:

R = ( t t ) - F* F - = 0 <311)

and

'312)The loading conditions are similar to those when the stress point lies on the

yield surface, except that when the stress increment, d<r,-y, directs outward from the

reference surface, R, it is defined as reloading. The described loading conditions are

shown in Figure 3.3.

3.2.3 H ardening Function, a^.

The hardening function may be expressed in terms of variables that are related to

the plastic deformations as follows:

Op. = a P. ( f ,6 >,6 '') (3.13)

where f a n d £y are the trajectories of total, deviatoric, and volumetric plastic

strains, respectively. In order to capture the non-dilative behavior of normally con

solidated clay, Wathugala (1990) proposed the following hardening function for the

£*-series models


2000

Yield surface

Reference surface

HC line

Current stress point Reloading

Neutral loading

Unloading

0 1000 2000

VTo^VTo,

Figure 3.3: The yield surface, F, and the reference surface, R, in the triaxial plane.


34

( f r + M f c ) * * ( 3 U )

where hi,h.2 ,h 3 and h4 are material parameters. The increments of and £v are

defined as

M d = (*& *& )* (3.15)

and

f ( 7 5 )*$- ^ ^ > 0 |div = \ V (3.16)

[ 0 for d e fr < 0 J

where defy = defy — |defray is the incremental deviatoric plastic strain tensor, and

def- is the incremental volumetric plastic strain due to virgin loading. Wathugala

(1990) carried out a parametric study about the effect of h3. It was found that this

hardening function can account for the behavior of both clays and sands.

3.2.4 Interpolation Functions

Evaluation of the plastic modulus, H *, is necessary to calculate the elastoplastic

constitutive stiffness and compliance tensors Cijki and Dijki- The plastic modulus

for virgin, H VL, is obtained from the consistency condition as:

f — ^h vl = Kaap.J L

where

r _ d ° b * ( Q Q \ a , d c i p ' f a k k ) . .

where < > are the McAuley’s brackets:


and 7i^t-y is the deviatoric part of the n - tensor.

The plastic moduli for non-virgin loading (unloading and reloading) ca n n o t be

obtained directly from the plasticity theory. Empirical functions are usually used

to evaluate them. Wathugala (1990) proposed the following simple interpolation

function to evaluate the reloading plastic modulus:

h r l = h v l + h v l Ti ^ r j ( 3 .2 0 )

where r i and r 2 are material parameters, and and Hj2L are virgin plastic moduli

at points I \ and J2 on the yield surface, as shown in Figure 3.3.

The intersection of the radial line passing through the current stress point and

yield surface locates point 7i, while the intersection of the hydrostatic compression

line and the yield surface represents the location of point / 2. The ratio (op^/ov) in

Equation 3.20 takes into account the distance effect from the current stress point to

the yield surface. When the stress point approaches the yield surface, av —* atp, and

—> H%L] which assures the smooth transition from reloading to virgin loading.

The presence of Hj7L in Equation 3.20 maintains the value of f f RL always greater

than zero, which is important since H* = 0 represents perfect plasticity and should

not occur inside the yield surface.

The most significant part of the plastic strains is developed during the reload

ing part of the cycle during cyclic loading. A small percentage of these strains is

developed during unloading. As a result, many investigators (Banerjee and Yousef,

1986; Dafalias and Hermann, 1986; Zienkiewicz et al., 1985; Wathugala and De-

sai, 1993) have successfully used the elastic unloading to predict the cyclic behavior

of clays. Despite the fact that an unloading plastic modulus, H UL, was proposed


36

by Wathugala (1990) for the HiSS £*-series models, elastic unloading behavior is

assumed to be satisfactory in the current study.

3.2.5 Potential Function, Q

The potential function was introduced in the plasticity theory to define the direction

of the plastic strain increments. In the case of the associative models, such as 5$,

the yield function itself is used as the potential function. Nonassodative models, on

the other hand, have separate potential functions. For example, the nonassodative

anisotropic <5£ model has a potential function with a translation rule to predict the

behavior of anisotropic days.

3.3 T he Proposed Interface H iS S -^ i M odel

Soil behavior at the soil-structure interface differs from its behavior in the far field.

High shear strains usually develop at the interface. Therefore, the analysis of soil-

structure interaction problems requires advanced constitutive laws to account for the

behavior of the material at the interface.

Due to pile installation and loading, severe shear deformation occurs at the soil-

pile interface. Experiments have showed that shear failure actually occurs inside

the soil near the soil-pile interface. During the Earth Technology Corp. (1986)

investigation, cyclic load tests were performed on pile segment models and afterwards

multiple shear-failure surfaces were observed in the day near soil-pile interface. As

a result of the high shear strain at the soil-pile interface, the effective stresses at

failure of a soil element at the interface are very low.

Field measurements of insitu stresses at the end of pile driving at Sabine Pass,

Texas, indicated that the actual stress state reached by the soil at the soil-pile

interface was very low. Furthermore, a laboratory study conducted on Boston Blue


37

Clay by Ahmad (1990) showed similar behavior (Figure 3.4). Laboratory direct

simple shear tests showed that the effective stress path of the soil almost reached a

point close to the origin. The low state of stress resulted from severe shear straining,

where shear strain reached almost 35%, and still pore water pressure increased and

normalized shear stress decreased, as shown in Figure 3.4b.

In this study, an HiSS constitutive model named Sh is developed in order to

capture soil behavior under severe shear deformation at the soil-pile interface. The

model is also applicable to the soil in the far field. This model is proposed based

on the more general nonassodative anisotropic 8 % model (Wathugala, 1990). The

anisotropic parameter is modified in this study so that a material point at the inter

face can reach a very low state of stress under severe shearing, as it does in reality.

Yield and potential functions for this model are the same as those of the 8 £ model,

(the yield function was described previously in this chapter).

3.3.1 P o ten tia l F u n c tio n , Q

The potential function, Q, is a small convex surface whose movement in the stress

space is based on a defined translation rule. The expression for Q surface is given

in terms of the modified stress tensor, &ij, and a constant value for the hardening

function, ocqo, as

« (» « ) = w - W - = 0 <3-21>

where

A , = ( £ ) + 7 ( £ ) (3.22)

F, = (1 - p S r)m (3.23)


Nor

mal

ized

sh

ear

stre

ss

Nor

mal

ized

sh

ear

stre

ss

38

0.4

0.3

0.2

0.1

0.00.0 0.2 0.4 0.6 0.8 1.0 1.2

Normalized effective vertical stress

(a)

0.3Shear stress

Pore pressure

0.2

0.1

0.00 10 20 30 40

Shear strain, %

(b)

Figure 3.4: Results of direct simple shear test on NC Boston Blue Clay,(a) Stress path; (b) Shear stress-strain and pore water pressure-strain curves (after Ahmad, 1990).


39

o-ij = °-ij - aij (3-24)

J \ = <Tkk (3.25)

J id — 2 (3.26)

h d = (3.27)

Si,- = o-ij - i <Tkk$ij (3.28)

the position and size of the potential surface in the stress space are defined by <ztJ-

and ago, respectively (other functions, variables, and parameters were defined in the

previous section).

3.3.2 Translation Rule

An efficient translation rule proposed for the model by Wathugala (1990) is

adopted herein. The isotropic potential function, Qi,0, is defined as the location

of Q tangential to the loading (reference) surface at the current stress point. The

isotropic translation tensor, atJ(t,0) , represents the position of the isotropic poten

tial surface, Qito- Figure 3.5 illustrates the isotropic surface and its location in the

Ji—\JJm space. The isotropic translation tensor can be expressed in terms of the

current stress, Cij, (Somasundaram and Desai, 1988) as

a ij{Uo) = <ra (3.29)

The translation tensor, atj, is defined as the deviation from atJ(,,0), and is given

by


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

30

(0£L

Loading surface (R or F) • Current stress point,

Isotropic potential surface, Q(lto) ♦ Location of a (hK)J

Locus o f possib le locations of a

20 -

10 -

100 200

J„ kPa

Figure 3.5: Location of the translation tensor, a^ , in the J,-VJ2D space (after Wathugala, 1990).

300

-uo

41

®»j — ®»j(mo) "I" dij (3.30)

the deviation tensor, dij, is separated into a deviatoric (/*»/), and isotropic (II) parts

as:

dij = II&/ + fiij (3.31)

The normalized 11(5) is related to the anisotropy parameter, a„, (Somasundaram

and Desai, 1988) by

5 = Aqan (3.32)

5 = i (3.33)Jim

J lm = ( — ) n_J Pa (3.34)\ a Q o J

where A q is the translation material parameter, and Jim is the maximum J\ on the

Q surface. The parameter a* is related to the amount of anisotropy, and is defined

as (Desai et al., 1986)

a ' ~ ! m J r (3 3 5 )

The anisotropic parameter is modified in the current study as follows

On — Aq\ Aq2 (3.36)

where Aqi and A q 2 are constants whose values range from 0 to 1 , and I \ and I 2 are

the first and second invariants of the strain tensor, respectively. Superscripts e and p

refer to elastic and plastic quantities, respectively. In order to simplify the model and


42

achieve numerical stability, the constants AqX and A q 2 are set 0 and 1, respectively.

Therefore, Equation 3.36 becomes

= ( 3 3 7 )

In order to find the direction of plastic strain increments, the translation tensor,

atj, and then the modified stress tensor, ay,-, should be determined. An illustration

of the translation rule proposed by Wathugala (1990) is shown in Figures 3.5 and

3.6. The position of the translation tensor, aij, can be located from the position of

the isotropic translation tensor, atJ(tJO). The isotropic potential function touches the

loading surface at the current stress point, r , in the stress space as shown in Figure

3.5. From Equations 3.30 and 3.31, the volumetric part of the translation tensor can

be written as

+ 3H (3.38)

In Figure 3.5, point p represents the center of Q(ito)- A shift of 311 distance on the

Ji-axis direction locates the center of Q in that direction. As a result, the translation

tensor, aij, lies on the plane A-A, as shown in Figure 3.5. This means that the shift

°f Otj(«o) in the Jx-axis direction is determined. The location of atJ- on the section

A-A in the octahedral plane is shown at point v in Figures 3.5 and 3.6. In addition,

the locus of possible locations for atJ- is also shown in Figure 3.6.

The direction of the deviatoric part of the translation tensor, tv, in Figure 3.6,

may depend on the strain history and the current stress. Wathugala (1990) defined

the direction as a combination of the deviatoric plastic and elastic strains

fij ~ ~^[^j + eij]

= -hey (3.39)


43

Locus of possible locations of

°2

Figure 3.6: Location of the translation tensor,a^; section A-A in the octahedral plane (after Wathugala, 1990).


44

where Cjy, e?-, and e?- are deviatoric total, plastic, and elastic strain tensors, respec

tively, and 6 is a positive scalar. Manipulation of the previous Equations will give

the translation tensor, a,y, (Wathugala, 1990) as

dij — 1 " ""1 ^ + + f—) " 3 S i* ~ 6e*\ O L Q o J \ a Q 0 j(3.40)

where

b =FbQF,

hD(3.41)

where F^q and F, are given in Equations 3.22 and 3.23, respectively. Alternatively,

the function F, may be evaluated from strain invariants as (Wathugala, 1990)

F. = ( l - ( 3 S r)m (3.42)

- y/T f h D /0 ^5 . - (3.43)

where I 2d and Izu are the second and third invariants of e^, respectively.

3.4 Evaluation of the M odel

This section presents the capabilities and characteristics of the proposed model.

3.4.1 Field Investiga tion on Sabine C lay

Successful accomplishment of this research required a well-documented case history

for the verification and evaluation of the proposed theoretical and numerical study.

The field and laboratory experiments carried out at a test site located 4 miles south

of Sabine Pass, Texas, satisfied this requirement. The Sabine Pass case study was

adopted in this study because it contains all necessary field and advanced testing

programs not found in other reviewed studies.


45

A considerable number of field and laboratory experiments have been conducted

on Sabine clay (Matlock and Tucker, 1961; Matlock and Bogard, 1973 and 1975;

Matlock and Holmquist, 1976; Bogard and Matlock, 1979; Grosch and Reese, 1980;

Earth Technology Corporation, 1986). The Earth Technology Corporation field in

vestigation (1986) was carried out on instrumented pile segment models installed

in Sabine clay. Measurements of total radial pressure, pore water pressure, shear

transfer, and relative pile-soil displacement during driving, consolidation, and static

and cyclic axial loading, were reported. In addition, an advanced laboratory testing

program (Katti, 1991) was performed on Sabine clay at the University of Arizona,

Tucson. Consolidated-undrained conventional triaxial compression (CU-CTC), hy

drostatic compression (HC), and truly triaxial tests were among the experiments

conducted.

Sabine clay is described as normally consolidated gray marine clay (Earth Tech

nology Corp., 1986). Table 3.1 presents some of its mineralogical, physical and

mechanical properties. These properties were reported by Earth Technology Corpo

ration (1986) and Bogard and Matlock (1979).

3.4.2 M aterial Param eters

In this study, the model was developed from the HiSS-tfJ model. Therefore,

evaluation of material parameters can be achieved following the same procedure

described for the HiSS ^‘-series models. Wathugala (1990) presented a detailed study

on this subject. Material parameters for Sabine clay as determined by Wathugala

(1990) are given in Table 3.2. Wathugala (1990) used these parameters to calibrate

the Sq model and to successfully predict soil behavior at Sabine Pass, Texas. The

same material parameters were used in this study to characterize Sabine clay behavior

through the 6 ^ model.


46

Table 3.1: Soil properties of normally consolidated Sabine clay (after Bogard and Matlock, 1979, and Earth Technology Corporation, 1986).

Mineralogy:

Montmorillonite 60%

Kaolinite 24%

Dlite 12%

Chlorite 4%

Physical properties:

Liquid limit, LL 100%

Plastic limit, PL 28%

Plasticity index, P I 72

Water content, w 73%

Submerged unit weight, 7 #tl$ 5.6kN/m3

Mechanical properties:

Undrained shear strength (UU-CTC) 24.4 kPa

Undrained shear strength (torvane) 35.9 kPa

Coefficient of consolidation, Cv 7 x 10- 9m2/sec

Preconsolidation pressure, Pa 95.8 kPa


4 7

Table 3.2: Material constants for Sabine clay (after Wathugala, 1990).

Elastic constants:

Shear modulus, G 4147 kPa

Bulk modulus, B 24536 kPa

Young’s modulus, E 11777 kPa

Poisson’s ratio, u 0.42

Basic plasticity parameters:

Ultimate parameter, 7 0.047

Ultimate parameter, /3 0.0

Ultimate parameter, m -0.5

Phase change parameter, n 2.4

Hardening parameters:

hi 0.0034

hi 0.78

hz 0

hA not applicable

Interpolation parameters (reloading):

T\ 500

r 2 2.4


48

3.4 .3 Im p le m en ta tio n o f th e M odel in to C o m p u te r P ro ced u res

The proposed model was implemented into the following:

1 . The computer program developed by Wathugala (1990) for the HiSS 6 ‘-series

models.

2. The computer code that utilizes the strain path method to simulate pile driving

effects on the surrounding soil.

3. The general purpose finite element computer program ABAQUS/Standard (Hi-

bbitt, Karlsson & Sorensen, Inc., 1995).

3 .4 .4 C apab ilities an d P ro p e rtie s of th e M odel

A major characteristic of the Vi model is that it allows the stress point to move

beyond the phase change line (or critical state line), as illustrated in Figure 3.7. In

the case of the ££ model, the failure occurs on the phase change line, which the stress

point cannot pass. In the proposed model, the degree of strain in the soil governs

the location of the stress point in the stress space. The failure point is controlled by

the parameter A q.

The capability of the model to predict failure beyond the phase change line de

pends on the magnitude of A q. In order to demonstrate this, the model was used to

predict different stress paths using Sabine clay parameters. The undrained conven

tional triaxial compression test was first simulated. Figure 3.8a shows the variation

of the predicted effective stress path during the test with different Aq values. The

soil underwent isotropic consolidation before the shearing; therefore, the stress path

moved from the origin to point A, in Figure 3.8a. Consequently, as a result of shear

ing, the stress point moved from point A to point C (failure state), when Aq, for

example, equaled —0.5. The corresponding stress-strain relationships for these tests


49

500

Ultimate line400

Phase change line

Yield surfacew 300

QCM"3

^ 200

Failure point predicted by VFailure point

predicted by s^'

100

20000 500 1000 1500 2500

J, kPa(a)

200

150

CDQ.-X.100aCN

0.00 0.04 0.08 0.12

^ * 2 D

(b)

Figure 3.7: (a) Comparison of predicted effective stress paths during undrained triaxial compression test using Sjj* and 8 * models; (b) The corresponding predicted stress- strain relationships.


50

500

Ultimate line400

Phase change line

« 300

Qr>4"5^ 200

100

0 500 1000 1500 2000 2500

J< kPa11

(a)200

150

(0CL

100aCM~5

0.00 0.04 0.08 0.12

VI2 D

(b)

Figure 3.8: (a) Variation of predicted effective stress path during undrained triaxial compression test with Aq; (b) The corresponding predicted stress-strain relationships.


51

are presented in Figure 3.8b. The soil samples experienced a peak shear stress value,

followed by the critical state wherein the strain increased while the shear stress re

mained constant. As the Aq value decreased, the model predicted a lower state of

stress.

The model was also used to predict the results of the simple shear test. Figure

3.9 presents the results of the simulated test. The predicted stress paths and stress-

strain relations are similar to those found in experiments by Ahmad (1990). Low

states of stress were measured during a direct simple shear tests conducted on Boston

Blue Clay (Ahmad, 1990), as shown in Figure 3.4. The model is therefore capable

of predicting a low stress state, as indicted by the preceding discussion of the results.

3.4.5 Calibration o f th e M odel

Calibration and evaluation of the proposed HiSS-tfJ,- model was needed in order to

check the capability of the model to predict soil behavior. Constitutive models are

usually calibrated and evaluated with respect to laboratory experiments. In the

field of geotechnical engineering, simple (conventional triaxial compression) and ad

vanced (true triaxial compression) laboratory experiments may be used, depending

on the nature of the proposed constitutive model. In the case of the proposed HiSS-

^2i model, the calibration required advanced laboratory experimental devices such

as the interface testing device (Rigby, 1996) and the direct simple shear (Ahmad,

1990). Unfortunately, such devices were not available to achieve this requirement.

Alternatively, the proposed HiSS-£Jt- model was calibrated on the basis of field ex

periments.

In order to calibrate the model and verify its capabilities, the field experiments

conducted at the Sabine site were utilized. Field measurements (Earth Technology

Corporation, 1986) of the total radial stress and pore water pressures were reported

during and after driving pile segment models. Figures 3.10 and 3.11 show the results


52200

150

co0.

- 100

500 1000 1500 2000 25000

J, kPa11

(a)

150

100(0Q.

aCN

0 «-0.0 0.1 0.2 0.3

(b)

Figure 3.9: (a) Variation of predicted effective stress path during simple shear test with Aq; (b) The corresponding stress-strain relationships.


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

roQ.

tntnJ=tnrau20)>

st=hi

150

Field measurements

125 &2i model prediction

100

25

i m i l l ijJ i i i i 11 il lm i j i i i 11 j i i i i

1.0E+0 1.0E+1 1.0E+2 1.0E+3

Time after pile driving, min

1.0E+4

Figure 3.10: Comparison o f predicted and measured effective radial stresses immediately after driving 1.72 in. diameter pile segment model into Sabine clay.

Oiu>

Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

125 t—i i 11 iii|----1—r i 111111-----1—i i 111 n|-----1—i i 111111-----1—i i 111111------1—n -

(0a.XL

tntn£tnraTJ2a>>••8£LU

75 -

50

25

Field measurements

^ 62,' model prediction

J — 1 1 11 m l -------1— L-i 1 m i l 1__1 1 11 m l 1 ........ 1 1 1 1 11 m l 1 11

1.0E-1 1.0E+0 1.0E+1 1.0E+2 1.0E+3


1 .OE+4

Figure 3.11: Comparison o f predicted and measured effective radial stresses immediately after driving 3 in. diameter pile segment model into Sabine clay.

55

of these measurements, at a depth of about 16 m below ground surface, after driving

1.72 in. and 3 in. diameter pile segment models. The 1.72 in. diameter pile segment

model field results were used to calibrate the model. Then back prediction was

performed using the field measurements of the experiments conducted on the 3 in.

diameter pile segment model.

The &2 i model was utilized to predict the stress field in the soil around the pile

immediately after driving. The strain path method was used to find the strain field

in the soil around the pile during driving. Different A q values were used in the predic

tion. The results are shown in Figure 3.12. The variations of the predicted effective

stress path with A q, for a soil element located 16 m below ground surface, in addition

to the stress-strain relations, are shown in the same Figure. The predicted effective

radial stress distribution around the pile, as it varied with Aq, is presented in Figure

3.13. As the value of Aq decreased, the model predicted a lower stress state. For

each Aq value the predicted effective radial stress at the soil-pile interface was com

pared with the field measurements. The value A q = —0.5 accurately estimated the

effective radial stress at the soil-pile interface. Therefore, the model was calibrated

for Sabine clay, where Aq equaled -0.5. Figure 3.10 shows the field measurements of

the effective radial stress immediately after driving as well as during consolidation of

the soil around the 1.72 in. diameter pile segment model. The predicted value of the

effective radial stress is shown in the same Figure. As demonstrated, the proposed

model successfully predicted the behavior of the soil in this experiment.

The variations of the distribution of the predicted pore water pressures with A q

are shown in Figure 3.14. These curves are associated with the effective radial stresses

presented in Figure 3.13. The model prediction of the pore water pressure at the

soil-pile interface was only 70% of the measured value. However, the effective stresses

were estimated accurately. The discrepancy in pore water pressure will appear in

the total stress and will not influence the prediction of the effective stresses. Since


56

20

K„ condition

CDQ.

QCM“D

A = 0.0

A = -0.3

A = -0.5

Aq= -0.7

A = -0.9

0 100 200 300

J-, , kPa

(a)

CDQ.

QCM

20

10

0.0 0.1 0.2

* 2 D

0.3 0.4

(b)

Figure 3.12: (a) Variation of effective stress path with Aq during pile driving as predicted by the model, using strain path method; (b) The corresponding predicted stress-strain relationships.


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

100 "i------------1-------- 1— i— i— i i i [

75roQ.</>'inSin.5 50TJS©>'■sit=UJ 25 A—A a

Q 0 - 0 ----0 O "0

□ □ B -------- □ □ □ □ H o

f r o O O 8 0

- v 00

- v a3

- V -0 5

s— v a7- 0

J i i i i i i

10 100

Normalized distance from pile wall (r/r0)

Figure 3.13: Predicted distribution o f effective radial stress for different Aq values immediately after pile driving, using the 5*2j model.

Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

(OCLX.e3(AWd)

£to£2on

250

225

200

175

15010 100

Normalized distance from pile wall (r/r„)

Figure 3.14: Predicted distribution o f pore water pressure for different Aq values immediately after pile driving, using the 8*2j model.

00

59

soil behavior depends on effective stresses, the model was considered satisfactory to

complete the proposed numerical study.

The model predictions of Sabine clay behavior were verified by simulating other

field experiments. Back prediction of the measured effective radial stress, at the

soil-pile interface around the 3 in. diameter pile segment model installed in Sabine

clay, is presented in Figure 3.11. The numerical results compare well with field

measurements, as shown in Figure 3.11. This demonstrates the capability of the

model to predict the behavior of the soil at Sabine Pass, Texas.


CHAPTER 4 SIMULATION OF PILE INSTALLATION

4.1 Introduction

The installation process severely remolds and disturbs the soil around the pile and

generates changes in the pore water pressures. These effects greatly influence pile

performance. Simulation of the pile driving process provides the initial conditions

for subsequent consolidation and loading. Therefore, it is necessary to adopt a reli

able method to simulate pile driving and then to obtain the deformation field in the

surrounding soil. The accurate estimation of these deformations results in appro

priate predictions of the stresses and pore water pressures. There are three major

methods that can be used to simulate pile driving: the cavity expansion method, the

strain path method, and the large-strain finite element method. Among the different

methods available to simulate pile driving, the strain path method, as proposed by

Baligh (1975, 1985), was successfully used to solve deep penetration problems such

as: simple pile (Baligh, 1975; 1984; 1985; 1986); cone penetrometers (Levadoux and

Baligh, 1980); open-ended piles (Chin and Baligh, 1983); soil samplers (Baligh et al.

1987); and closed-ended pile segment models (Wathugala, 1990). This method was

used to simulate pile driving in the current study.

This chapter presents the aspects of simulation of pile driving using the strain

path method. Regarding the present study, the strain field around the closed-ended

pile (simple pile) was obtained using formulas, while a numerical scheme was used

to develop a computer code for the open-ended pile.

60


61

4.2 M ethods o f Pile Driving Simulation

The behavior of piles driven into saturated clay is governed by the effects of driv

ing, subsequent consolidation, and loading on the clay-water system. In order to

understand the behavior of the pile, these effects may be studied using the most

reliable and practical methods. Theoretical and numerical studies, using different

methods, have been conducted to simulate these effects. Cavity expansion methods,

or CEMs, are the most popular ones used to simulate pile driving. The simplicity

of CEMs (due to the one-dimensional nature of the cavity expansion) has enabled

investigators to obtain analytical and numerical solutions (Butterfield and Banerjee,

1970; Carter et al., 1979).

However, the CEM cannot capture the two dimensional nature of the pile driving

problem. Because the actual deformations (due to driving) near the top and the tip

of the pile do not occur radially, as the cylindrical CEM postulates. In addition, the

strain path assumed in the CEM is not realistic and can lead to errors because soils,

in general, are strain path dependent. While cavity expansion methods are helpful

in developing a qualitative understanding of the problem, they are not a complete

solution.

Exact simulation of pile driving is difficult. The usual procedure would be to

start from an initial state of stress and pore water pressures and then to simulate the

process of driving as the pile is pushed into the ground. Desai (1978) proposed using

a large-strain finite element formulation for this purpose. Kiousis et al. (1988) used

a similar approach to solve the cone penetration problem, utilizing the large—strain

finite element method and the Cap model (DiMaggio and Sandler, 1971). The Cyber

205 supercomputer took 20 cpu hours for 20 mm advancement of the cone penetrom

eter. Consequently, with available computing power, it would be very expensive to

exactly simulate pile driving, which needs severed meters of advancement.


62

Evidence from Experiments (Rourk, 1961; Vesic, 1963; Robinsky and Morrison,

1964; Szechy, 1968; Randolph, et al., 1979) confirmed that the deformations caused

by the penetration of a rigid indenter axe similar in different soils, even though

the penetration resistance can be drastically different. This implies that the pile

installation is a strain controlled problem and that the deformations associated with

it are not very sensitive to soil behavior (Levadoux and Baligh, 1980). Based on

these observations, Baligh and co-workers (1975, 1984, 1985, and 1986) developed

the Strain Path Method (SPM), to obtain solutions for different deep penetration

problems including pile driving. Their method is based on getting a kinematically

admissible deformation field for the particular problem, by using analytical solutions

of uniform inviscid flow around sinks and sources. By distributing different sources

of variable strength in the uniform flow, deformation fields were obtained around: (a)

the simple pile (Baligh 1975, 1984, 1985, and 1986), (b) cone penetrometers (Baligh

and Levadoux, 1980), (c) open ended piles (Chin and Baligh, 1983), and (d) soil

samplers (Baligh et al., 1987). The strain path at each point can be calculated from

the deformation field. Then the effective stress path corresponding to the strain

path can be determined from the constitutive equations. Total stresses and pore

water pressures can be obtained by solving the equilibrium equations. It should be

noted here that the final stresses and pore pressures will be exact if and only if the

estimated deformation fields are identical to those experienced in the actual problem.

Tumay et al. (1985) and Acar and Tumay (1986) proposed a strain path method

based on a semi-analytical technique to obtain deformation patterns around cone

penetrometers (this method uses the conformal mapping rather thcin sources and

sinks, as in the strain path method).


63

4.3 Strain Path M ethod

The strain path method is an approximate analytical technique that utilizes the so

lution for steady irrotational flow of an inviscid incompressible fluid to solve deep

penetration problems. The basic idea of this approach is to obtain a kinematically

admissible deformation field for the particular problem by superimposing flow due

to different sources and sinks into a uniform steady flow coming from infinity. Com

bining a uniform steady flow with a flow resulting from a simple source yields the

solution for the simple pile, while replacing the simple source with a ring source

produces the solution for the open-ended pile.

4.3.1 Limitations

The strain path method simulates deep, steady, quasi-static, two-dimensional, undrained

penetration of axisymmetric piles in saturated (incompressible) homogeneous isotropic

clay initially subjected to an isotropic state of stress. These conditions cannot be met

in the case of piles driven into A"0-consolidated soils. Moreover, the method neglects

the surface roughness of the pile wall. This implies that the predicted deformation

field will only be approximate and therefore stresses and pore pressures estimated by

this method will also be approximate. However, experimental observations (Rourk,

1961; and Randolph et al., 1979) have suggested that the deformation field predicted

through these approximations may be considered the best available solution for deep

penetration problems.

4.3.2 C losed-Ended P iles

The simple pile approach (Baligh, 1975) was proposed to investigate the installation

effects of the simple solid (closed-ended) pile on the surrounding soil. The solution is

based on the principles and assumptions of the strain path method. Superimposing


64

the flow resulting from a simple source (a point of outward radial flow) and a uniform

steady flow, provides the formulation for the simple pile problem.

Considering a cylindrical coordinate system with radial axis r and vertical axis z,

the stream function for a uniform steady flow of velocity U is given by (Thompson,

1976):

rp = Ur2 (4.1)

the stream function corresponding to a simple source emitting a fluid of volume equal

to 47rm per unit time (where m is the strength of the source) is given by (Thompson,

1976):

ip = m cos (p = m ------------j- (4-2)(r2 + z2)?

superposition of the stream functions, in Equations 4.1 and 4.2, yields:

ip = ~ ~ U t 2 + mcos <p (4-3)

the dividing stream line equation can be written as:

(^ ) = ^ [1 + cos^ (4-4)

this equation gives the geometry of a blunt-nosed cylindrical body whose diameter

at z = oo or (<p = 0) is 2R which is known as the simple pile. The geometry of the

simple pile is shown in Figure 4.1.

The path of a soil particle initially located at a radial distance r„ from the center

line of the pile can be obtained from (Baligh, 1985a):

<«>where R is the radius of the pile shaft.

R eproduced with perm ission of the copyright owner. Further reproduction prohibited without perm ission.

65

Stream line

► r

Stream line not drawn to scale

Figure 4.1: Geometry o f the simple pile.


66

The semi-analytical solution for strains obtained by Baligh (1985) is only ap

proximate, because it is based on the simplifying assumption that the stream lines

are vertical. This implies that this solution is not exact and is only valid in the far

field, where stream lines are nearly parallel to the pile axis. Teh and Houlsby (1987)

considered the actual curved profile of the stream line and derived the following exact

analytical expressions for strains around the simple pile:

(4.6)

= -Fl(rf) - j(l - 3B*)fi(*) (4.7)

(4.8)

- (2B y /B 2 - l) tan -1 cot

sin(2$)

(4.9)

where err, ezz, egg, and eTZ are the radial, vertical, tangential, and shear strain

components, respectively, and

= 3(1 + cos (<£))[ ©

2 3 cos (<f>)(4.10)

F2(4) = b fl + i (^) (1 + cos(0)) (4.11)

B = 2(s ) ! + i <412>The modified strain field Equations 4.6, 4.7, 4.8, and 4.9 (Teh and Houlsby, 1987)

were used in this study to estimate the strain field in the soil around the closed-ended


67

(or simple) pile. Figure 4.2 shows the predicted strain field around the simple pile for

soil elements located initially at radial distance ra = R and ra = 2R, respectively. In

this figure, the amount of strain is almost negligible when the soil element is located

at a vertical distance more than 5R in front of the tip of the pile. Then, as the soil

element approaches the tip of the pile, higher strains are induced by driving. The

magnitude of the radial and tangential strains increases as the soil element passes

the tip and then remains constant at a vertical distance of about 5R behind the tip

of the pile. The vertical strain vanishes at almost the same vertical distance, while

shear strain decreases to a very low value then remains constant. The variation with

radial distance of strains in the soil around the pile at the end of driving is presented

in Figure 4.3. High strains are developed at the pile-soil interface. The amount of

strain decreases with the increase in the radial distance away from the pile until it

becomes negligible in the far field.

4.3.3 O p en -E n d ed Piles

The mechanism of open-ended (cookie cutter or totally unplugged) pile driving was

investigated, via the strain path method, by Chin and Baligh (1983). Simulation

of an open-ended pile installation was performed using the method of sources and

sinks, by combining the flow of a ring source with a uniform steady flow. In the strain

path method, the pile is considered an “ideal” open-ended pile if it is simulated by

superimposing one ring source on a uniform steady flow. Different pile geometries can

be obtained by considering different numbers and arrangements of the ring sources in

the uniform flow. Using the cylindrical coordinate system, the formulation presented

herein describes the penetration of axisymmetric “ideal” open-ended pile.


68

5.0

tr

8 2 5 -ceg]35T 3

ro0.0 -

®>TJ®NtO j c _§ 2 5 Oz

-5.0- 0.2-0.4 0.20.0 0.4

Strain, %(a)

5.0

CE15g 2.5ctoCOT315

o.o -0)>TJ0)N1 -2.5 -

r J R = 2.0

oz

-5.0- 0.2-0.4 0.20.0 0.4

Strain, %

(b)

Figure 4.2: Strain field around simple pile for (a) Soil element located initially at r0=R, (b) Soil element located initially at r0=2R.


69

CM

in

©ocra"wT>raTJ5

T 30)NraEo

04

o£I!N.£H.c3S00c

■3o

' i

<D.2•3UHE

&e*3oco3

r2*CC/3scnrOw300

in ino

oo

tno'

% ‘UIBJJS


70

4.3.3.1 R ing Source

The ring source is a distribution of constant strength sources along a circle. The

velocity components of a ring source of radius R placed in the plane z — 0 and

emitting an incompressible fluid at a rate of volume V per unit time in an infinite

medium are (Kuchemann and Weber, 1953):

V

4 * 2 v j z 2 + (r + R ) 2

v° = V 2 z-E(k)

(4.13)

(4.14)‘ 4* 2 yjz2 + (r + R ) 2 [z2 + (r - fl)2]

where v° is the radial velocity, v° is the vertical velocity (in the cylindrical coordinate

system), and k is the particle position parameter which is given by

4 tRh z 2 + (r + R ) 2 (4‘15)

the functions K {k ) and E(k) are complete elliptical integrals of the first and second

kind, respectively, and are given by:

K M = nJ o V r^ fc s in ^ a

da (4.16)

E(k) = f 3 y jl — k sin2 a da (417)Jo

the complete elliptical integrals can be evaluated through the following polynomial

approximation (Abramowitz and Stegun, 1964):

K{m) = + axm x + a2 m \ + a3 m \ + (4-18)

(=r)+ [&o + bxm x + b3 m \ + b3m l + In f — 1

+ e(m) | e(m) |< 2 x 10 - 8


71

where m i = 1 — m, and 0 < m < 1; and

a0 = 1.38629436112

a\ = 0.09666344259

a2 = 0.03590092383

a3 = 0.03742563713

a4 = 0.01451196212

and

b0 = 0.5

&i = 0.12498593597

b2 = 0.06880248576

b3 = 0.03328355346

64 = 0.00441787012

E(m ) = £l + a-[TTi\ + a2 m \ + 03771 + a4m ij

+

+ e(m)

where

= 0.44325141463

a2 = 0.06260601220

a3 = 0.04757383546

a4 = 0.01736506451

+ b2 m \ + b3 vn\ + 64m^j In

| e(m) |< 2 x 10'

fcx = 0.24998368310

b2 = 0.09200180037

63 = 0.04069697526

64 = 0.00526449639

(4.19)

4.3.3.2 Superposing a Uniform Steady Flow on the Source—Ring Flow

The solution for the open-ended pile is achieved by superposing a uniform steady

flow of velocity U on the flow of a ring source. As a result, an annular body of infinite

length is obtained, which is the geometry of the open-ended pile. Figure 4.4 shows a

diagram for the geometry of the open-ended pile. In a cylindrical coordinate system,

the velocity components of a particle moving in the combined flow are


B = 2R

soilInnerOutersoil

Particle path

Particle path not drawn to scale

Figure 4.4: Geomerty of the open-ended pile.


73

vr = v° and vz = v° + U (4.20)

where vr and vz are the radial and vertical velocity components, respectively, and v°

and v° are the radial and vertical velocity components of the ring source as given by

Equations 4.13 and 4.14, respectively.

The relation between the aspect ratio of an open-ended pile, B / t (where B is

the diameter and t is the wall thickness of the open-ended pile), and the discharged

volume of the source fluid, V, can be obtained from the continuity as follows:

V = t U ( R 2 - R]) (4.21)

where R and Ri axe the outer and inner radii of the open-ended pile far behind the

tip, respectively. Defining the wall thickness t as t = R — f?,-, Equation 4.21 can be

approximated as :

V ~ 2trRtU, 4 < 1 (4.22)K

or

B B2irU— ~ ——— where B = 2R (4-23)

this approximation is acceptable since most of the open-ended piles in the engineering

practice have an aspect ratio of 17 < B / t < 44 (Azzouz and Baligh, 1984).

4.3.3.3 S tra in F ield

The coordinates of a material point Mi a t time t with initial coordinates r° and z°

far enough in front of the tip of the pile are given by

ri = r° + [ Vr(ri>zi)dT (4.24)Jo


74

z\ = z\ + f v z ( t J , zT)dr (4.25)Jo

where vr and vz axe the radial and axial velocity components in the cylindrical

coordinate system, respectively. These components are given by Equation 4.20. The

tangential velocity vg vanishes because of axial symmetry.

The rate of deformation tensor, Dij, is given as follows

where ut- are the velocity components. The following components of the rate of

deformation tensor, D ^, during axisymmetric penetration in an infinite medium are

derived from Equation 4.26.

n - dVr n - dVz n — Vf n _ 1 ( dVr a o7^r r ~ dr ' z z ~ d z ' _ r ’ r z ~ 2 [ d z + d r ) ( }

where vr and vz are the radial and vertical velocity components, respectively.

The natural strain increment, detJ-, is introduced to determine the strain field in

the soil around the pile during driving

£ij = J deij = J Dijdt (4.28)

the integration is conducted by following a particle on its path along the stream

line. During undrained penetration, the resulting strain components must satisfy

zero volume change condition, which means:

£rr + + £se — 0 (4-29)

The estimation of deformations, strain increments, and strains in the soil during

open-ended pile penetration is performed numerically, due to the complexity of the


75

ring source formulation. A numerical procedure proposed by Levadoux and Baligh

(1980) for investigating the deformations and strains around the cone penetrometer

was adopted by Chin and Baligh (1983) for the open-ended piles. This method

was used, with modifications on the integration method, in this study to develop

a computer code to simulate the driving of open-ended piles. The following is a

general description of the main steps of this method:

1. The path of a soil particle during driving is obtained by numerical integration

of Equations 4.24 and 4.25. The radial and axial velocity components are

given by Equation 4.20. Small time increments are used with the mid-point

integration method.

2. During the particle motion along its path, at each time increment the com

ponents of the rate of the deformation tensor are calculated from Equation

4.27.

3. Natural strain increments are obtained by the numerical integration of Equa

tion 4.28 at each time increment. Strain components at time t are obtained by

adding the strain increments.

The strain field in the soil during the driving of open-ended piles was obtained

using the described numerical procedure. The effect of the penetration velocity on

the predicted strain field was investigated. Simulation of pile driving was performed

with different penetration velocities. Results indicated that the penetration velocity,

U, did not affect the predicted strain field, even though the rate of deformation of the

soil around the pile was influenced. Figure 4.5 shows the strain field determined by

the strain path method resulting from the driving of open-ended piles with different

penetration velocities.

The predicted strain field for a soil element located, at the end of driving, at

distance 77 = 1.5fZ outside the wall, is shown in Figure 4.6a. The strain field for


Nor

mal

ized

ve

rtic

al d

ista

nce,

z/R

76

4 .0

2.0

0.0

Results match for(a) U = 0.02 m/s(b) U = 0.20 m/s

- 2.0

-4.0-0.04 - 0.02 0.00 0.02 0.04

Strains, %

Figure 4.5: The effect of pile driving velocity, U, on the strain field of the surrounding soil (soil element is located at radial distance r = 1.5 R).


Reproduced

with perm

ission of the

copyright ow

ner. Further


without

permission.

a:■aofocain‘o8■e<1)>~o0)NID§oz

4

2

0

■2

-0.04 - 0.02 0.00 0.02 0.04

O'

ofoca</)8€a>>■oa>Nra§o

4

2

0

■2

•rz

-0.04 - 0.02 0.00 0.02 0.04

Strain

(a) r,=1.5R

Strain

(b) r,=15t

Figure 4.6: Strain field predicted by SPM due to an open-ended pile driving; (a) Soil element outside the pile located finally at radial distance rf=1.5R; (b) Soil element inside the pile located finally at radial distance r,^15t.

-j

78

this element is similar to the strain field obtained around the simple pile. Figure

4.6b presents the strain field for a soil element located inside the wall at a distance

77 = 15f from the center of the open-ended pile at the end of pile driving.

The strain paths for two soil elements located, at the end of driving, in the

immediate vicinity of the pile wall are presented in Figure 4.7. The strain paths are

presented in the deviatoric f?t—space where Ei, Et and E3 are defined as:

Ei = ezz

Et = ^ (e V r - )

E3 = ~^2 ,£rz (4.30)

the strain path moved along E\ for the triaxial test, along Et for the pressuremeter

test and along E 3 for the direct simple shear test. As shown in Figure 4.7, the soil

elements were located at 77 = r — 1.2£ inside and 77 = r + 1 outside the wall of the

pile. These elements were subjected to severe shearing and strain reversals. The

outside element, for example, experienced about 60% strain along the Ez-zxis (i.e.

the direct simple shear mode).

4.4 Numerical Simulation o f Pile Driving

In this study, the general procedure described by Wathugala (1990) is modified and

used to obtain the distribution of effective stresses, total stresses, pore water pres

sures, and hardening parameters in the soil, as follows:

1. The strain path method provides the strain path for a soil particle during pile

driving. Determination of the strain path due to pile driving is performed at

all Gauss points in the finite element mesh.


CJII

liT

- 0.1

o.o

0.1

0.2

0.4

= r-1.2t

= r+t0.3 -

s t / /

-0.2 -0.1 0.0

E2 = (Err* %)A/3

0.1

CO-ycu_CMII

COHI

0.6

0.4

0.2

0.0

rf =r-1.2t

rf = r+t

- 0.2- 0.1 0.0- 0.2 0.1

B,= (V e«t)/V3

Figure 4.7: Strain paths, in the deviatoric space, o f soil elements located at the immediate vicinity outside and inside the wall during open-ended pile driving; (a) E, vs E2, (b) E3 vs E2.


80

2. The corresponding effective stress path for each point is determined by inte

grating the constitutive equations along the given strain path. The proposed

H iSS-^i model through the ‘strain to stress’ algorithm is utilized to fulfill this

purpose.

3. The pore water pressures, effective stresses, and hardening parameters calcu

lated in step 2 are supplied to the finite element program ABAQUS/Standard

as input data. Distributions of pore water pressures, total stresses, effective

stresses, and hardening parameters that simultaneously satisfy equilibrium

equations, strain compatibility, and constitutive equations are computed by

ABAQUS/Standard. These distributions are calculated under undrained con

ditions, because pile driving is assumed to occur under these conditions. The

results at the end of the pile driving simulation are used as initial conditions

for the subsequent consolidation stage.

This procedure was applied to simulate the driving process for closed-ended and

open-ended piles. The effective stress path during driving for a soil element at the

pile-soil interface is presented in Figure 4.8. Point A in the Figure represents the

AT0-condition before pile driving. As the pile penetrated the soil, the state of stress

decreased until point B was reached at the end of driving. The distribution of effective

stresses at the end of driving is depicted in Figure 4.9. As shown in the Figure, a

low state of stress was predicted for the soil at the pile-soil interface as a result of

severe shearing. ff0-conditions were estimated for the soil in the far field, where

the strains developed due to driving were negligible. The corresponding predicted

pore water pressure distribution is shown in Figure 4.10. High pore water pressure

was predicted at the pile-soil interface, while hydrostatic pore water pressure was

estimated in the far held.


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without

permission.

Stress path during pile driving

Yield surface

Phase change line

20raCL

0 50 150100 200 250 300

J, , kPa

Figure 4.8: The predicted effective stress path for a soil element at the pile-soil interface during pile driving.

00

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permission.

150

125(UCL-sc 100</)'w0)5 75u>0)I 50

LLI or

'oo

■e—o o o ooooocojooaMiiixx)— e—o o o o o o o o o

1 10 100Normalized distance from pile wall (r/r„)

Figure 4.9: Predicted distribution of effective stresses immediately after pile driving.

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without

permission.

(OCL

8>3atatfl>

0)ro£a>oQ.

250

225

200

175

1 _ J L j . i X J 1 I I I J .150 x x j i i i

10 100


Figure 4.10: Predicted distribution of pore water pressure immediately after pile driving.

OOu>

CHAPTER 5VERIFICATION: PILE DRIVING,

SUBSEQUENT CONSOLIDATION AND PILELOAD TESTS

5.1 Introduction

In this study, a general method for predicting pile setup is proposed and verified. The

solution is based on a consistent formulation during the different life stages of a pile.

Utilized to achieve the objectives of this research are: the HiSS modeling approach

(Desai et al., 1986; Wathugala, 1990; Wathugala and Desai, 1993), the strain path

method (Baligh, 1975, 1984 and 1985; Levadoux and Baligh, 1980; Chin and Baligh,

1983), and the coupled theory of the nonlinear porous media formulated in the finite

element program ABAQUS/Standard (Hibbitt, Karlsson & Sorensen, INC., 1995).

The soil behavior was characterized by the proposed HiSS-^,- model (discussed

in detail in Chapter 3). The strain path method was used to simulate the pile driving

process. The conditions at the end of driving were used in the finite element program

ABAQUS/Standard to start the consolidation phase. This program uses the cou

pled theory of the nonlinear porous media to solve different geotechnical engineering

problems including consolidation. During the consolidation phase, different pile load

tests were simulated.

It is necessary to verify and evaluate the proposed method based on existing field

experiments. The verification scheme is important in order to check the validity of

the numerical solution and to asses the possibility of using the method in future

applications.

This chapter presents the verification of the proposed method. Field experiments

performed on a 1.72 in. (4.37 cm) diameter pile segment model installed at Sabine

84


85

Pass, Texas, were used for this purpose. The verification included the pile driving

process, subsequent consolidation, pile load tests conducted during consolidation,

and the final cyclic pile load test.

5.2 Field Tests at Sabine Site

The field experiments conducted by the Earth Technology Corp. (1986) on Sabine

clay were employed to verify and evaluate the current theoretical and numerical

study. These experiments were performed at a depth of 50 ft (15.24 m), using two

instrumented pile segment models of 1.72 in. (4.37 cm) and 3 (7.62 cm) in. in

diameter. Changeable cutting shoes with different wall thicknesses were attached

to these models in order to conduct measurements on open-ended piles. Table 5.1

presents the dimensions of these pile segment models. These models were equipped to

measure simultaneously the total radial pressure, pore pressure, shear transfer, and

relative pile-soil displacement, during driving, consolidation, and static and cyclic

load tests.

The testing program was initiated: (a) to examine the effects of the diameter/wall

thickness ratio, D /w t , on the rate of consolidation and setup; (b) to study the effects

of the diameter, D, on the rate of consolidation and setup; and (c) to investigate

the role of the effective radial pressure and pore water pressure with regard to the

axial shear transfer capacity. This well-documented study provided high quality

experimental data in the area of axial pile-soil interaction and therefore was used

for the verification and evaluation scheme.

5.2.1 T est Procedures

The test procedure sequence for each experiment is described in Figure 5.1. One

of the main objectives of these experiments was to study the evolution of the static


86

Table 5.1: Dimensions of the pile segment models used in the field experiments.

Diameter (D), in. (cm) Wall thickness (wt), in. (cm) D/wt

1.72 (4.37) - 2 (closed-ended)

3.0 (7.62) - 2 (closed-ended)

3.0 (7.62) 0.125 (0.318) 24 (open-ended)

3.0 (7.62) 0.065 (0.165) 46 (open-ended)


87

Setting up the loading system at the top of the borehole

Drilling and casing the bore holes up to the test depth

Lowering the pile segment model into the borehole

Attachment of the data acquisition system to the pile segment model

Performing the final two-way cyclic load tests at the end of consolidation

Conducting static tension and compression tests during consolidation

nstallation of the models by pushing the X-probe or driving the 3 in. diameter pile segment model

Continuously measuring the pore pressure, the shear transfer, the total radial pressure, and the relative soil-pile displacement

Figure 5.1: The test procedure sequence for the field experiments.


88

shear transfer capacity during consolidation. Therefore, it was decided to perform

load tests at degrees of consolidation of 20, 40, 60, 80 and 90 percent and more.

However, the load tests were conducted at degrees of consolidation close to the

predetermined ones.

5.2.2 P ile Segm ent M odels

Two pile segment models with changeable cutting shoes were used in Sabine field

experiments: the 1.72 in. diameter pile segment model (also known as the X-probe)

and the 3 in. diameter pile segment model. The 1.72 in. diameter pile segment

model is shown in Figure 5.2. This model has a length of 56.5 in. (1.44 m). It

is equipped with a total pressure transducer, a pore pressure transducer, a shear-

sensing element (friction sleeve), and a displacement transducer. Figure 5.3 shows

the 3 in. diameter pile segment model. This probe is instrumented with: load cells

(for friction measurements), total and pore pressure transducers, and DC-LVDT (for

measurements of the relative pile-soil displacement). Detailed information about

these models and the test setup was presented in the studies conducted by Earth

Technology Corp. (1986) and Wathugala (1990).

5.2.3 R esults of Field Experim ents

Presented herein are the results of the field experiments conducted on the 1.72 in.

diameter pile segment model. The measurements taken during the pile driving are

not available. The measured total radial and pore water pressures during the con

solidation phase are shown in Figure 5.4. The corresponding calculated effective

radial stress variation is also presented in the same figure. The breaks in the plots

correspond to the time at which pile load tests were performed. It can be noted that

the static load tests have only minor effects on the consolidation phase (Earth Tech

nology Corp., 1986). Pile load tests were performed during consolidation in order to


89

INSTRUMENT CABLE

STANDARD CONE ROD CONNECTION

LOAD CELL

TOTAL PRESSURE CELL

PORE PRESSURE CELL

DISPLACEMENT TRANSDUCER

SOIL ANCHOR

Figure 5.2: The 1.72 in. diameter pile segment model (after Earth Technology Corp., 1986).


90

O R IL L RO D ADA PTER

Ca b l e j u n c t i o n

SPACER

LOAO CELL

SPACER

HOU SING FO R TO TA L A N D PORE PRESSU R E

TRAN SD U CERS

SPACER

LOAO C ELL

O O L V D T HOUSING

i n t e r c h a n g e a b l e CUTTIN G SHOE

• SEE O E T A 1L A

> SEE / D E T A IL b

; SEE / D E T A IL A

ITlUUf; tAZtZD E T A IL A

Ulf VI MtXU |LOAO CELL FOR

FRICTION MEASUREMENT

' SEE D ETA IL C

D E T A IL B

T O T A L P R E S S U R E L O A D C E L L A N D P O R E P R E S S U R E

T R A N S D U C E R H O U S IN G

t 0 E T A 1 L C

I D C -L V D T F O RI M E A S U R E M E N T O F

/ R E L A T IV E D ISPL A C E M E N Te r r t iM C s n o c

Figure 5.3: The 3 in. diameter pile segment model (after Earth Technology Corp., 1986).


Effe

ctiv

e ra

dial

pre

ssur

e, k

Pa

Radi

a to

tal

and

pore

pr

essu

re,

kPa

91

400

300

200 -

Total radial pressure

Pore water pressure

10010 100 1000 100001

Time after driving, min

(a) Total radial and pore water pressures.

200

100 -

10 1000 100001 100


(b) Effective radial pressure.

Figure 5.4: Variations of the measured pressures during soil consolidation around the 1.72 in. diameter pile segment model (after Earth Technology Corp., 1986).


92

investigate the increase of bearing capacity with time (the pile setup). The results

of the measured shear transfer-displacement response for different tension pile load

tests are shown in Figure 5.5. Other results will be presented and compared with

the numerical solutions later.

5.3 T he Finite Element Program A B A Q U S/Standard

The general purpose finite element program ABAQUS/Standard (called ABAQUS in

this report) was used in the current study to perform the numerical simulation of the

consolidation phase and the pile load tests. The results of the numerical simulation

were processed through the post processing program ABAQUS/Post.

5.3.1 A nalysis o f Porous M edia

The analysis of porous media is treated in ABAQUS by the conventional approach.

The porous medium is considered a multi-phase material and the effective stress

principle is used to describe its behavior. The theory of deformation of porous media

as presented in the ABAQUS Theory Manual (Hibbitt, Karlsson & Sorensen, INC.,

1995) is summarized below. ABAQUS notations and sign conventions are used here.

The effective stress, a, acting at a point in the saturated porous medium, is given

by:

<r = <r + x u j i (5.1)

where a is the total stress, uw is the wetting liquid pressure (pore pressure), and

X is a factor that depends on the saturation and surface tension of the liquid-solid

system. For saturated systems, x = 1-0. In ABAQUS, tensile stress components are

stored as positive, which explains the sign in Equation 5.1.


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permission.

S . 20uT(ACsI—ro<u■Cw

Degree of consolidation

U = 22%

U = 60%

U = 87%

U = 98%

0.0E+0 3.0E-4 6.0E-4 9.0E-4 1.2E-3 1.5E-3

Displacement, m

Figure 5.5: Results o f the field static tension tests conducted on the 1.72 in. diameter pile segment model (after Earth Technology Corp., 1986).

v Ou>

94

The constitutive response of the porous medium includes the simple bulk elasticity

relationships for the pore fluid and the solid grains. Moreover, the material skeleton

behavior is determined by the constitutive theory in which the effective stress, <f , is

defined as a function of the strain history, the temperature, and the state dependent

variables.

The behavior of the porous medium is governed by the equilibrium and the con

tinuity equations. For the solid phase of the material, the stress equilibrium is

expressed by the virtual work principle for the volume V in its current configuration

at time t as:

J ( a - xtt*I) : SedV = J t ,8vdS+ j f.8vdV + J snpiofT.SvdV (5-2) v s v v

where 5e = sym(d8v/dx) is the virtual rate of deformation, a is the Cauchy effective

stress, 6 v is a virtual velocity field, t are surface tractions per unit area, f are body

forces per unit volume (excluding fluid weight), pw is the fluid density, and g is

the gravity acceleration. The Lagrangian formulation is used to discretize the solid

phase equation with the displacements as the nodal variables. In order to model the

porous medium, the continuity equation is required for the fluid phase. The equation

is obtained by equating the rate of increase in fluid volume stored at a point with

the rate of volume of fluid flowing into the point within the time increment:

/T ( I I = ~ f ^ r 37m -v todS (5-3)\ v P°w J s P°m

where v„ is the average seepage velocity (fluid velocity relative to the solid phase),

n is the outward normal to the surface S, s is the degree of saturation, and n is

the porosity. This continuity equation is normalized by the reference density of the

fluid p°w. The backward Euler approximation is used for the time integration of the


95

continuity equation. For further details, the reader is referred to Section 2.6 in the

ABAQUS Theory Manual (Hibbitt, Karlsson & Sorensen, INC., 1995).

5.3.2 Im plem entation o f th e HiSS—£|t- M odel in A B A Q U S

The proposed EiSS-S^ model was implemented in the finite element program ABAQUS,

which allows the use of user defined subroutines. The HiSS- model was first in

corporated in the ‘strain to stress’ algorithm developed by Wathugala (1990) for the

HiSS-£* series models. This algorithm uses the incremental theory of plasticity to

predict the stress increment corresponding to a given strain increment during vir

gin loading, unloading, and reloading. The hardening parameters are also updated

during the incremental calculations.

The user defined subroutine, UMAT, was then used to implement the ‘strain

to stress’ algorithm into ABAQUS. UMAT is developed in ABAQUS to incorporate

any user defined material behavior into the finite element. This subroutine exchanges

information between the ‘strain to stress’ algorithm and ABAQUS. The information

includes the element number, Gauss point number, step number, iteration number,

time, total strains, incremented strains, predicted stresses, and updated hardening

parameters.

The initial conditions needed to be given to the finite element program before

starting the analysis. This was achieved through the user defined subroutines SIGINI

and SDVINI. The user subroutine, SIGINI, defines the initial stress field. The initial

stresses are calculated at each Gauss point in the finite element mesh. The subroutine

SDVINI is called to define the initial solution dependent state variable fields only

in the beginning of the analyses. The variable fields include the strain trajectories,

plastic strains, initial hardening function, and the flag that defines the virgin loading,

unloading, and reloading.


96

The numerical simulation was performed in this study through a single set of user

subroutines. First, the pile driving process was simulated. The effective stresses and

hardening parameters were determined during pile driving at all Gauss points in the

finite element mesh. This was achieved through the ‘strain to stress’ algorithm used

to integrate the constitutive equations. The integration was performed along the

strain path for the Gauss point during pile driving. The strain field was predicted by

the strain path method. At the end of driving, ABAQUS performed globed iterations

to calculate the equilibrated effective stress and pore water pressure distributions.

The effective stresses, pore water pressures, and hardening parameters calculated at

end of driving were supplied to ABAQUS as initial conditions for the consolidation

phase. During the consolidation, each pile load test was simulated at a time equal

to the actual time of conducting the field test. Figure 5.6 shows a diagram for the

numerical simulation procedure.

5.3.3 F in ite Elem ent Mesh

The element library in ABAQUS contains a wide range of elements which provide

the geometric modeling capability for various applications. Due to the symmetry

around the pile axis, the axisymmetric solid element CAX8RP was selected to model

the geometry of the problem. This element is one of a variety of elements provided for

modeling coupled pore fluid diffusion/stress analysis problems. The nodal variables

for this element are the displacements and the pore pressure. The element CAX8RP

uses second-order (quadratic) interpolation for the geometry and displacements. The

pore pressure is interpolated linearly from the comer nodes. Reduced integration

scheme is recommended by Wathugala (1990) to avoid unstable prediction of pore

water pressures in this type of problems. This element is generally described as an

8-node biquadratic displacement, bilinear pore pressure, with reduced integration

scheme. The element CAX8RP and its characteristics are shown in Figure 5.7.


97

Open-ended pile i Call KOPENDRIV!

Closed-ended pile Call KSPILE

Run ABAQUS

Call pile driving subroutines KINSITU (main)

Initial conditions before pile driving Call SIGINI Call SDVINI

_ Initial conditions at | the end of pile driving

; ABAQUS-UMAT

Stresses, pore pressures, hardening parameter and state dependent variables at the end of pile

j driving (initial conditions for consolidation)

Pile load tests including the final cyclic load test

Consolidation stresses, pore pressures, hardening parameters

(initial conditions for pile load test)

Figure 5.6: Diagram for the use of UMAT in the finite element analysis.


98

4 7 3

8

21 5

Uz▲

p • --------------------► U r

• Node

+ Gauss point

up Radial displacement

uz, Vertical displacement

p, Pore water pressure (at nodes 1, 2, 3, and 4)

Figure 5.7: ABAQUS axisymmetric solid element CAX8RP used in the finite element analysis.


99

The finite element mesh used to simulate the consolidation and pile load tests for

the 1.72 in diameter pile segment model is shown in Figure 5.8. The mesh consists

of 578 elements and 1837 nodes. The analysis was started using a coarse finite

element mesh. This mesh led to numerical oscillations in pore water pressures. The

mesh was then refined until these numerical oscillations disappeared. The mesh was

generated so that small size elements were used around the pile. This was necessary

for accurately capturing the response in the surrounding soil where laxge deformations

were developed. Moreover, the fine mesh arrangement leads to numerical stability

and faster convergence.

The H iSS-^i model was proposed in this study to characterize the soil behavior

at the interface. As presented in Chapter 3, this model has the capability to predict

the stress state beyond the phase change line (critical state line). Such a stress

condition results from the high shear deformations at the soil-pile interface. On the

other hand, the HiSS-£j model predicts the state of failure at the phase change line.

Therefore, up to a certain degree of straining, both models predict almost similar

stress conditions before the phase change line. This is illustrated in Figure 5.9. The

strain field during pile driving for soil elements (A, B, C and D) located at different

radial distances from the pile surface was determined. Then the HiSS-j^ model was

used to predict the corresponding stress path for these elements, as shown in the

figure. The stress points at the end of driving passed the phase change line when soil

elements were located at radial distances 1.5R and 5R (elements C and D). On the

other hand, the stress point stopped before the phase change line for the element at

6.5R from the pile center line (element B). The far field element A remained under

AT0-condition both before and after driving.

The predictions of the HiSS-tf ,- and models were similar when the elements

were located at a radial distance of more than about 6.5R from the center line of

the pile. Therefore, it was expected that the interface model BSSS-fJ,- would be used


fi> <5 T

< 5

< s

Element. 8-node ABAQUS-CAX8RP 578 elements 1837 nodes

Figure 5.8: The finite element mesh used to discretize the soil domain around the 1.72 in. diameter pile segment model.


Pile Ground surface

(a) The location of the elements.

CDO.

□CM

30

Yield surface

Phase change line

Radial distance = 1.5R

Radial distance = 5R20

Radial distance = 6.5R

10

00 100 200 300

J1 , kPa

(b) The stress paths.

Figure 5.9: Variation with radial distance of predicted effective stress paths during pile driving for soil elements located 16 m below the ground surface.


102

for the soil around the pile up to about 6.5R radial distance, and the HiSS-£g model

would be used for the soil beyond this distance. However, the difference between the

predictions of the HiSS-tfJ,- and Sq models is negligible before the phase change line

is reached. Therefore, the whole soil domain around the pile was modeled using the

interface HiSS- f t model.

5.4 P ile Driving

As described in Figure 5.1, the 1.72 in. diameter pile segment model was driven at

the bottom of the bore hole. The soil from the ground surface down to the bottom of

the bore hole was assumed to be at AT0—condition. The soil below the bottom of the

bore hole was affected by pile driving. The strain path method was used to estimate

the strain field in the soil around the pile during driving. The 1.72 in diameter pile

segment model is a closed-ended pile and therefore the pile driving was simulated

utilizing the simple pile approach (Baligh, 1984). Simulation of pile driving was

presented in detail in Section 4.3.

The soil element located at the soil-pile interface at a depth of about 16 m was

used for verifying and evaluating the numerical results. This depth corresponded to

the location of the radial and pore water pressure transducers in the field experiments.

The predicted effective stress path for the soil element during pile driving is shown

in Figure 5.10. Point A in the figure represents the AT0-condition before pile driving.

As the pile penetrated the soil, the stress point moved along the path until it reached

point B at the end of pile driving.

Shown in Figure 5.11 are the variations of the predicted effective stresses with

the normalized radial distance immediately after pile driving. Examination of this

figure indicates that pile driving caused a significant decrease in the effective stresses


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30


Yield surface

Phase change line

nCL

0 50 100 150 200 250 300

J, , kPa

Figure 5.10: Predicted effective stress path for a soil element at the soil-pile interface during driving the 1.72 in. diameter pile segment model.

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permission.

(0Q.yU)in

in0)>'•8*=l b

150

125

100 'oo

75

50

25

0

1 10 100Normalized distance from pile wall (r/r0)

Figure 5.11: Predicted distribution of effective stresses at 16 m below the ground surface, immediately after driving 1.72 in. diameter pile segment model.

o

105

near the pile, up to a radial distance of about ten times the pile diameter. The pre

dicted effective radial stress, for example, decreased from 69 kPa before pile driving

to 17 kPa immediately after pile driving. The effective stresses in the far field re

mained under iif0-condition. A comparison of the predicted and measured effective

radial stresses immediately after driving is shown in Figure 5.12. The prediction is

consistent with the field measurements.

The proposed HiSS-^,- model successfully predicted the low stress conditions

immediately after driving. Field experiments (Earth Technology Corp., 1986; Azzouz

and Lutz, 1986; Azzouz and Morrison, 1988; Bogard and Matlock, 1990) showed that

low effective radial stresses were obtained immediately after pile driving. In a study

conducted in Empire, Louisiana, Azzouz and Lutz (1986) found that the effective

radial stresses acting on the pile shaft during installation were smaller than the

initial insitu effective radial stresses. The solution obtained in the current study is

consistent with the findings of these field studies. On the other hand, theoretical

and numerical studies (Randolph et al., 1979; Heydinger and O’Neill, 1986; Chopra

and Dargush, 1992) presented contrary results, showing that the predicted effective

radial stresses increased immediately after driving.

The predicted pore water pressure distribution and the maximum measured value

at the pile wall are shown in Figure 5.13. The maximum pore water pressure pre

dicted at the soil-pile interface was about 70% of the measured value. The maximum

excess pore water pressure was under predicted by a factor of 0.5. Similar results

were obtained by Baligh and Levadoux (1980) for cone penetrometers in clays, where

the excess pore water pressure was under estimated by a factor of 0.5. Pore water

pressure was also under predicted by a factor of 0.3 during the pile behavior study

by Wathugala (1990), where the HiSS-^J model was used to model the entire soil

domain.


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150

Field data

125 k =1.0E-8 m/sec

k=1.0E-9 m/sec

k =6.25E-10 m/sec

ro■5(0k . Immediately after

driving0)>•■83=at

25

1.0E+0 1.0E+1 1.0E+2 1.0E+3 1.0E+4


Figure 5.12: Comparison of predicted and measured effective radial stresses during soil consolidation around the 1.72 in. diameter pile segment model.

OC\

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ner. Further


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permission.

400

£3tntna>q . 200L .

B(0£2o

100

Field measurement

Finite element prediction

1 10 100


Figure 5.13: Predicted distribution of pore water pressure immediately after driving 1.72 in. pile segment model.

o

108

For given boundary conditions, many possible stress fields satisfy the equilib

rium requirements. In the simulation of pile driving, the analysis began with an

approximate solution for effective stresses, strains, and displacements in order to

calculate the stress field under equilibrium. The results of this analysis showed that

the calculated effective stresses are consistent with the measured values. Therefore,

the predicted effective stresses obtained by the constitutive model are the correct

effective stress fields. In such condition, many combinations of the to tal stresses and

pore water pressures can produce one effective stress field. In this method, the pore

water pressures were under predicted, and therefore the total stresses were also under

predicted.

The preceding results and discussion indicate that even though the pore water

pressure is under predicted, the effective radial stress at the end of driving is in good

agreement with the field measurement. The comparison is shown in Figure 5.12 (the

value correspond ing to coefficient of permeability k = 1 x 10~9 m/sec).

5.5 Consolidation and P ile Load Tests

The finite element simulation was conducted using the nonlinear transient analysis in

ABAQUS. The initial conditions required to start the simulation of the consolidation

phase were obtained from the end of pile driving stage. These conditions included

the pore water pressure and effective stresses distributions which were already under

equilibrium, in addition to the hardening parameters and other state dependent

variables.

During the simulation of the consolidation phase, different pile load tests (ten

sion, compression, and cyclic) were numerically simulated according to the sequence

presented in Figure 5.14. Pile load tests were conducted at degrees of consolidation

of 22, 40, 60, 87, 97, 98, and 99 percent.


109

Perform static tension test at U=60%





Perform static compression test at U=97%

Perform final two-way cyclic load tests at U=99%

Install 1.72 in. full displacement pile segment model

Perform static tension test at degree o f consolidation, U=22%

Figure 5.14: Sequence o f the load tests conducted during soil consolidation around the 1.72 in. diameter pile segment model.


110

The coefficient of permeability, k, was evaluated for Sabine clay, based on three

one-dimensional consolidation tests performed by the Earth Technology Corp. (1986).

In the present study, k was calculated as 6.25 xlO -10 m/sec and was used in the sim

ulation of the consolidation. The simulation was repeated using two other values for

the coefficient of permeability: 1.0xlO -8 and 1.0 xlO -9 m/sec. The eifect of k on

the rate of dissipation of the excess pore water pressure is illustrated in Figure 5.15.

The results indicate that when k equals 1.0 xlO -9 cm/sec, the predictions and field

measurements are consistent. Moreover, this value for k predicts the effective radial

stress distribution during the consolidation very well, as shown in Figure 5.12.

Using the o n e -d im ensional consolidation test, the evaluation of the coefficient of

permeability is only approximate. In the field, more reliable techniques usually axe

utilized to evaluate the coefficient of permeability. Because of the uncertainty in

evaluating k and because the difference between the calculated and corrected k is

relatively small, it is believed that the correction is within acceptable deviation. In

this study, therefore, the k value of 1.0xlO -9 m/sec was used in the simulation. This

value was also used in Chapters 6 and 7 for the closed-ended and open-ended piles,

where the results were found to be consistent with field measurements.

During the consolidation phase, the excess pore water pressures were gradually

dissipated. This dissipation was associated with the increase in the effective stresses.

The predicted pore water pressure distributions at different time intervals during

consolidation are shown in Figure 5.16. The corresponding changes in the effective

radial stresses are presented in Figure 5.17. At the end of consolidation, the excess

pore water pressures were fully dissipated and the m axim um increase in the effec

tive stresses was reached. Figure 5.18 shows the predicted variation of the effective

stresses at the completion of the consolidation. Examination of the figure indicates

that the effective radial stress increased from 17 kPa immediately after driving to

97 kPa at the end of consolidation. Moreover, the effective radial stress increased


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1.25 i n ui|£3tntn£ 1.00 -

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k =1.0E-8 m/sec

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1.0E-2 1.0E-1 1.0E+0 1.0E+1 1.0E+2 1.0E+3 1.0E+4 1.0E+5


Figure 5.15: Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 1.72 in. diameter pile segment model.

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250 1 sec

1370 s e c

6920 secraQ. 13600 se c

22524200 se ca>

L .3Wtn 39700 sec

63100 sec

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1501 10 100


Figure S. 16: Predicted dissipation of pore water pressures during soil consolidation around the 1.72 in. diameter pile segment model.

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HI

100 r - i - T

80

60

40

20

t--------- 1------1— i—i—r*T"T"

Q—©—©-----©—©■■ Q~^

I I _l____ I___I__I_I_I I I

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— B — 13800 sec

— K— 24200 sec

— ♦ — 39700 sec

-----A---- - 63100 sec

- 900000 sec

j ____ i___i__i_i_i i i1 10

Normalized distance from pile wall (r/rc)

100

Figure 5.17: Predicted increase in effective radial stresses during the consolidation of the soil around the 1.72 in. diameter pile segment m odel.

OJ

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125

z z

(0!* 100</)'w&(Aa>>■*3uI 751U

501 10 100

Normalized distance from pile wall (r/rc)

Figure 5.18: Predicted distribution o f effective stresses at a depth o f 16 m at the end of soil consolidation around the 1.72 in. diameter pile segment model.

115

from 69 kPa before pile driving to 97 kPa at the completion of consolidation. The

effective stresses in the far field were not affected during the consolidation. The ef

fective stress path for the soil element at the soil-pile interface during consolidation

is shown in Figure 5.19. Stress point C represents the stress condition at the end of

consolidation.

The simulation of pile load tests was performed during the consolidation phase.

These tests consisted of tension pullout tests, a compression test, and the final cyclic

pile load test. The initial conditions necessary to simulate a pile load test were

obtained from the consolidation phase at a time corresponding to the degree of

consolidation at which the load test was conducted. These conditions include stresses,

pore water pressures, and hardening parameters.

The measured displacement-time history was applied to simulate the pile load

tests. In Figures 5.20, 5.21, 5.22, 5.23, 5.24, and 5.25, the finite element results

for pile load tests # 1, # 2, # 3, # 4, # 5, and # 6 are compared with the field

measurements. The shear transfer represents the average values along the shaft of the

1.72 in. diameter pile segment model. Comparisons of the predicted and measured

variations of shear transfer during these tests indicate that the simulation captured

the measured behavior very well. Predicted variations of the shear transfer with the

displacement match the measurements for pile load tests # 3, # 4 and # 6 . Other

tests showed slight differences between the predictions and measurements, but the

shear transfer values at failure were well predicted.

The predicted variations of the effective radial stresses for these tests compared

very well with the field measurements. During pile load test # 3, the estimated and

measured variations of effective radial stress contradicted each other. However, the

predicted values showed good agreement with field measurements at the beginning

and end of the test.


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Stress path during consolidation

Yield surface

Phase change line20(0CL

oN—i

0 50 100 150 200 250 300

Jt , KPa

Figure 5.19: Predicted effective stress path for a soil element at the soil-pile interface during pile driving and consolidation o f soil around the 1.72 in. diameter pile segment model.

Shea

r tr

ansf

er,

kPa

Shea

r tr

ansf

er,

kPa

117

20

10

0 Degree of consolidation, U = 22%

Field pile load test

Finite element simulation

-10 ----0.0E+0 3.0E-4 4.0E-4 5.0E-41.0E-4 2.0E-4

Displacement, m

(a) Shear transfer vs displacement

20

10

\ -


Reid pile load test

Rnite element simulation

-10150 20050 1000

Time, sec

(b) Shear transfer vs time.

Figure 5.20: Comparison of predicted and measured response during pile load test #1 conducted on the 1.72 in. diameter pile segment model.

(figure con'd)


Effe

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kPa

118

40

30

20

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00 50 100 150 200

Time, sec

(c) Effective radial stress vs time.

400

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1000 50 100 150 200

Time, sec

(d) Pore water pressure vs time.

Degree of consolidation, U = 22%



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Shea

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Displacement, m

5.0E-4 5.5E-4


20

15

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5 Field pile load test


025 50 750 100 125

Time, s e c



(figure con'd)


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-

40 —

Degree of consolidation, U = 40%30 —

---------- Field pile load test—

---------- Finite element simulation -

20 1 , 1 1 . , 1 , 1

0 20 40 60 80 100 120

Time, se c


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12140


Reid pile load test


30

20

10

0

4.0E-40.0E+0 8.0E-4 1.2E-3

Displacement, m


40




50 750 25 100

Time, sec



(figure con'd)


Pore

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ater

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Pa

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kPa

122

100

75

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Degree of consolidation, U = 60%25 Field pile load test


00 20 40 60 80 100

Time, se c


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250 -_r

200 -

150 -Degree of consolidation, U = 60%



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Time, sec

80 100



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r tr

ansf

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kPa

Shea

r tr

ansf

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kPa

123

40




20

10

0

4.0E-42.0E-40.0E+0

Displacement, m


40


30 Reid pile load test


20

10

0

300200 400 5000 100

Time, sec



(figure con'd)


124

120

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- Degree of consolidation, U = 87% —

---------- Field pile load test

---------- Finite element simulation

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100 200 300

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Shea

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125

30

20

10

Degree of consolidation, U =97%



-10 ----0.0E+0 4.5E-3 6.0E-33.0E-31.5E-3

Displacement, m

(a) Shear transfer vs displacement.

30

20

10




-10200 250 30050 100 1500

Time, sec



(figure con'd)


Pore

w

ater

pre

ssur

e, K

Pa

Effe

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dial

str

ess,

kPa

126

300



Finite element simulation200

100 2000 300

Time, se c

(c) Effective radial stress vs tune.

r—200

100




00 100 200 300

Time, se c

(d) Pore water pressure.


Shea

r tr

ansf

er,

KPa

Shea

r tr

ansf

er,

kPa

30


Reid pile load test

Rnite element simulation15

0

-15

-30 ------3.2E-3 -2.4E-3 -1.6E-3 -8.0E-4 0.0E+0

Displacement, m


30


Field pile load test15


0

-15

-30400 6000 200 800 1000

Time, s e c



(figure con'd)


Pore

w

ater

pre

ssur

e, k

Pa

Effe

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dial

str

ess,

kPa

128

100

75

50

Degree of consolidation, U = 97%25 Field pile load test


0400 600 800 10002000

Time, sec


200 -

'--------------- j

100 -




200 400 600

Time, sec

800 1000



129

Since the pore water pressure was underestimated at the end of pile driving, the

discrepancy occurred in the first few pile load tests. Comparisons of the measured

and predicted changes in the pore water pressure during pile load tests # 4, # 5,

and # 6 indicate that the results are not consistent with each other. Generally, the

pore water pressure is not predicted well in this study.

Pile load test # 7 was numerically simulated at a degree of consolidation of

98 percent. Loading-unloading cycles were performed during the test. The finite

element results, as compared to the field measurements, are depicted in Figure 5.26.

The predicted variations of the shear transfer during the test are in good agreement

with field measurements, as shown in Figures 5.26a and 5.26b. The predictions

accurately capture the behavior during the loading-unloading cycles. Variations of

predicted effective radial stresses and pore water pressures are consistent with field

measurements (Figures 5.26c and 5.26d).

The final two-way cyclic load test was conducted near the end of consolidation.

The test consisted of five tension-compression cycles. During all cycles, the failure

occurred before the direction of loading was reversed. A comparison of the predicted

and measured shear transfer variations with displacement is shown in Figure 5.27.

The simulation predicted the response during the first cycle very well. During the

other cycles, the estimated shear transfer was slightly lower than that of the field

measurements. The variation of shear transfer during the test is presented in Figure

5.28. The predicted and measured response are in good agreement. The variations

with time of the effective radial stress and pore water pressure are shown in Figures

5.29a and 5.29b, respectively. The variations in the predicted effective radial stress

and pore water pressure during the test were relatively smaller than those in the

measured values. However, the finite element results at the beginning and end of the

test aue in good agreement with the field measurements.


Shea

r tr

ansf

er,

kPa

Shea

r tr

ansf

er,

kPa

130

40

20

0

-20Degree of consolidation, U = 98%

Reid pile load test


-40

-60 -----0.0E+0 1.5E-35.0E-4 1.0E-3 2.0E-3

Displacement, m


40

20

0

-20Degree of consolidation, U = 98%

Field pile load test-40


-60200 400 600 8000 1000

Time, sec



(figure con'd)


Pore

w

ater

pre

ssur

e, k

Pa

Effe

ctiv

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dial

str

ess,

kPa

131

100

75

" \

50




4000 200 600 800

Time, se c

(a) Effective radial stress vs time.

250

200




1000 200 400 600 800

Time, se c

(b) Pore water pressure vs time.


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40

20

coQ.Mil> owc£L .

W -20

-40

-1.5E-3 -1.0E-3 -5.0E-4 0.0E+0 5.0E-4 1.0E-3 1.5E-3

Displacement, m

Figure 5.27, Comparison of predicted and measured response (shear transfer vs displacement) during the final cyclic pile load test conducted on the 1.72 in. diameter pile segment model.

T



J i I i I ■ I ■ I

u>K>

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40

20

-V

(0c£&_CO0)JC0)

-20


Finite element simulation-40

2000 400 600 800 1000 1200

Time, sec

Figure 5.28: Comparison of predicted and measured response (shear transfer vs time) during the final cyclic pile load test conducted on the 1.72 in. diameter pile segment model.

U>U>

134

200


Finite element simulation(0CL

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4000 800 1200

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Finite element simulationoCL

4000 800 1200

Time, sec


Figure 5.29: Comparison of predicted and measured response during the final cyclic pile load test conducted on the 1.72 in. diameter pile segment model.


135

5.6 P ile Setup

The main objective of the current research was to predict pile setup, or the increase

of bearing capacity with time after pile driving. In order to accomplish this task,

the finite element simulation of pile load tests during consolidation was conducted.

Comparisons of the predicted and measured behavior of the pile in these tests showed

that the finite element results are in good agreement with field measurements.

For some load tests, the finite element simulation accurately predicted the mea

sured shear transfer at failure. For each simulated pile load test, the estimated peak

and residual shear transfer values were compared with those obtained in the field.

The results are presented in Figure 5.30. The comparisons showed that the predicted

increase in the shear transfer with time after pile driving are consistent with the field

measurements.

It should be noted that at the beginning of consolidation, the finite element

method overestimated the pile setup. The reason could be that the constitutive

model is not calibrated using overconsolidated test data. The finite element method

underpredicted pile setup at the end of soil consolidation.


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(0Q.i_ra(ACroL -(0a>JC(0

30

20

Field test, peak10

Field test, residual

— A -— Prediction, peak

— ©•— Prediction, residual

0 — 1.0E+3 1.0E+4 1.0E+5

Time after pile driving, sec

Figure 5.30; Comparison of predicted and measured increase in soil-pile shear transfer at failure (skin friction) for pile load tests performed at different time periods after driving.

U>O s

CHAPTER 6 NUMERICAL SIMULATION OF THE

BEHAVIOR OF CLOSED-ENDED PILES

6.1 Introduction

In the previous chapter, a proposed numerical method for investigating the increase in

bearing capacity, with time, of friction piles driven into saturated clay, was verified

with respect to field experiments. The proposed HiSS-^i model was calibrated

and successfully predicted the soil behavior at Sabine Pass, Texas. Then numerical

simulations of pile driving, subsequent consolidation, and load tests were conducted.

The numerical results were verified with respect to the field experiments conducted

on the 1.72 in. diameter pile segment model. During the verification stage, material

parameters, soil conditions, and numerical stability and convergence parameters for

Sabine clay were identified. It was essential to utilize this numerical environment to

evaluate another set of field experiments. The back prediction scheme would support

and assure the validity of the method.

The main objective of the current research was to study the increase in the

shaft capacity, with time, of friction piles driven into saturated clay. Therefore,

finite element numerical simulations of the different load tests conducted during

consolidation on the 3 in. diameter pile segment model, were employed to achieve

this goal. These load tests provided the qualitative and quantitative estimates of the

time increase in shaft capacity (or pile setup).

This chapter presents the results of the numerical simulation of the various life

stages of a closed-ended pile segment model driven into saturated clay.

137


138

6.2 Field Experim ents

Field experiments on a 3 in. diameter pile segment model were conducted during the

Earth Technology Corp. investigation (1986). The experiments consisted of measur

ing the total radial pressure, pore water pressure, shear transfer, and relative soil-pile

displacement, during driving, consolidation, and load testing. The 3 in. diameter

pile segment model was installed according to the procedure described in Figure 5.1

(refer to Section 5.2). Figure 6.1 presents the variations in the measured total radial

and pore water pressures as well as the change in the corresponding effective radial

pressure during consolidation. During consolidation, a series of pile load (tension

and cyclic) tests were conducted at different time intervals corresponding to degrees

of consolidation of 18, 37, 92, and 100 percent. Figure 6.2 shows the variation of the

shear transfer with the displacement for the conducted tension pile load tests. The

final two-way cyclic pile load test was performed at the completion of consolidation.

Figure 6.3 displays the sequence of these tests.

6.3 Finite Elem ent M esh

The 8-node axisymmetric solid element CAX8RP was used in the finite element

analysis (the element and its properties were described in Section 5.3.3).

The soil mass was discretized according to the finite element mesh presented in

Figure 6.4. The mesh consisted of 578 elements and 1837 nodes. The described mesh

(Figure 6.4) was used for the finite element simulation of consolidation and pile load

tests. The soil mass surrounding the pile was modeled as the HiSS-tfJi material.


Effe

ctiv

e ra

dial

pre

ssur

e, k

Pa

Tota

| ra

dia|

and

po

re

pres

sura

kP

a139

400

300

200 -Total radial pressure

Pore water pressure

10010 100 10001 10000


(a) Total radial and pore water pressures.

100

50

10 1001 1000 10000



Figure 6.1: Variation of the measured pressures during soil consolidation around the 3 in. diameter pile segment model (after Earth Technology Corp., 1986).


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30

20

CD0 .

&tncs

\

I—(0a)szCO U = 18%

U - 37%

U = 92%-10 U = 100%

5.0E-40.0E+0 1.0E-3 1.5E-3 2.0E-3 2.5E-3

Displacement, m

Figure 6.2: Results of static tension tests conducted on the 3 in. diameter pile segment model (after Earth Technology Corp., 1986).

O

141


erform static tension test at U=100

Perform static tension test a t U=37%

Install 3 in. full displacement pile segment model

Perform final two-way cyclic load test at U=100%


Figure 6.3: Sequence o f the load tests conducted during soil consolidation around the 3 in. diameter pile segment model.


142

s >

§ >

S > ,

T

< 5

aA

Element &-node ABAQUS-CAX8RP 578 elements 1837 nodes

Figure 6.4: The finite element mesh used to discretize the soil domain around the 3 in. diameter pile segment model.


143

6.4 Pile Driving

The 3 in. diameter pile segment model was driven according to the procedure de

scribed in Figure 5.1. The simple pile approach (Baligh, 1984), developed for deep

penetration of closed-ended piles, was utilized to obtain the strain field in the soil

during pile driving (details of this subject were described in Section 4.3). Simulation

of pile driving was performed in a manner similar to the simulation of the 1.72 in

diameter pile segment model (Section 5.4).

Throughout this chapter (unless otherwise indicated), the numerical results for

the soil element located at the soil-pile interface at 16 m below ground surface, are

compared with field measurements. This depth (16 m) corresponds to the approxi

mate location of the total and pore water pressure transducers during testing.

The predicted effective stress path for the soil element at the soil-pile interface

during driving is shown in Figure 6.5. Points A and B represent the initial K a-

condition before, and the condition immediately after pile driving, respectively. In

this study, simulation of pile driving reduced the effective mean stress (J i/3 ) as well

as the second invariant of the deviatoric stress tensor {Jzd)- At the end of driving,

the effective stresses for the soil around the pile were substantially decreased. On

the other hand, the effective stresses for the soil in the far field remained under K a-

conditions. Examination of Figure 6.6 indicates that the decrease in the predicted

effective stresses around the pile extended to about ten times the pile diameter. The

effective radial stress decreased from 70 kPa before, to 18 kPa immediately after pile

driving. The prediction of low effective radial stress at the end of pile installation

is consistent with the field measurements (Figure 6.7). The predicted pore water

pressure distribution at the end of pile driving is depicted in Figure 6 .8 . Shown

in the same figure is the measured pore water pressure, at the soil-pile interface,

immediately after pile driving. The numerical simulation predicted about 66% of


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permission.

30


Yield surface

Phase change line20

toQ.

Q04—)

0 100 200 300

J , , kPa

Figure 6.5: Predicted effective stress path for a soil element at the soil-pile interface during driving 3 in. diameter pile segment model.

Effe

ctiv

e st

ress

, K

Pa145

150

125

100

75

50

25

0

Figure 6 .6 : Predicted distribution of effective stresses at 16 m below the ground surface immediately after driving the 3 in. diameter pile segment model.

A A A A A .

G—G C O QQQQG8E

1 10 100

Normalized distance from pile wall (r/rj


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125

Field data

k =1.0E-8 m/sec(0 100 Q.

k=1.0E-9 m/secminSinTo

k =6.25E-10 m/sec

T32 After driving50a>>'Sa=LU

i 11ill i i m il___

1.0E+41.0E-1 1.0E+0 1.0E+1 1.0E+2 1.0E+3


Figure 6.7: Comparison of predicted and measured effective radial stresses after driving, and during soil consolidation around the 3 in. diameter pile segment model.

-u0\

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300


Field measurementtoQ.

O. 200

OCL

1001 10 100


Figure 6.8: Predicted distribution of pore water pressure immediately after driving 3 in. diameter pile segment model.

4*."4

148

the maximum pore water pressure measured at the pile wall. The maximum excess

pore water pressure was underestimated by a factor of 0.41. The maximum pore

water pressure was underestimated by a factor of 0.5 in the case of the 1.72 in.

diameter pile segment model (Chapter 5).

6.5 Consolidation and Pile Load Tests

The finite element simulation of the subsequent consolidation and load tests was

conducted according to the same procedure used for the X-probe in Chapter 5. Fig

ure 6.3 shows the sequence of simulating the pile load tests during the consolidation

phase.

The predicted and measured effective radial stresses during consolidation are com

pared in Figure 6.7. The finite element results indicate that the simulation predicted

the field behavior very well. The analysis was performed with the calibrated value

of the coefficient of permeability (k = 1 x 10-9 m/sec). This comparison supports

the argument presented in Section 5.2 regarding the coefficient of permeability.

During consolidation, the effective stresses gradually increased as the excess pore

water pressures dissipated. A comparison of the predicted and measured normalized

excess pore water pressures is presented in Figure 6.9. The numerical results are in

good agreement with the field measurements, and the pattern of the measured dis

sipation is accurately estimated. Most of the excess pore water pressures dissipated

within three hours after driving. The predicted pore water pressure distributions

at different time periods during consolidation are shown in Figure 6.10. The corre

sponding estimated increase in the effective radial stresses is depicted in Figure 6.11.

At the end of consolidation, the maximum increase in the effective stresses devel

oped upon complete dissipation of the excess pore water pressures (Figure 6.12). The


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permission.

1.25

3inw£ 1.00Q.

0.75

o0.50

inina>oa> 0.25

T 3<UN0.00

Field data

k=1.0E-8 m/sec

k=1.0E-9 m/sec(0EU . k =6.25E-10 m/secoz xuJ_

1.0E+4-0.25

1.0E+0 1.0E+1 1.0E+2 1.0E+3


Figure 6.9: Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 3 in. diameter pile segment model.

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fljQ_

e3(0Wa>

£CO

£oOu

250

11 sec

74 sec225

290 sec

1154 sec

4610 sec200

38434 se c

178434 sec

900000 sec175

O Q OOOQi

1501 10 100

Normalized distance from pile wall (r/rQ)

Figure 6.10. Predicted distribution o f pore water pressures during consolidation o f the soil around the 3 in. diameter pile segment model.

L/iO

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permission.

n>CLXLin(/>0)-bV)raTJ2a>>'■8i»=ID

100

80

60 11 sec

74 sec 290 sec

1154 sec40

4610 sec

38434 sec

178434 se c 900000 se c

20

01 10 100


Figure 6.11: Predicted increase in the distribution of effective radial stresses during consolidation of the soil around the 3 in. pile segment model.

Effe

ctiv

e st

ress

, kP

a152

125

100

75

50

251 10 100

Normalized distance from pile wall (r/ro)

Figure 6.12: Predicted distribution of effective stresses at a depth of 16 m at the end of soil consolidation around the 3 in. diameter pile segment model.


153

predicted effective radial stress increased from 18 kPa immediately after driving to

98 kPa at the end of consolidation. The final estimated effective stresses near the pile

increased when compared with the pre-driving conditions (ifo-conditions). For ex

ample, the predicted effective radial stress increased from 70 kPa before pile driving

(iT0-condition), to 98 kPa at the end of consolidation. In the far field, on the other

hand, the effective stresses of the soil remained unchanged during consolidation.

The predicted effective stress path during consolidation for the soil element at the

soil-pile interface is shown in Figure 6.13. Point C represents the stress condition at

the end of consolidation.

Four static tension pile load tests were simulated during the consolidation phase.

In addition, a simulation of the final two-way cyclic load test was conducted at the

end of consolidation (Figure 6.3).

The first three static pile load tests were performed in the tension mode. These

tests were simulated at degrees of consolidation of 18, 37, 92, and 100 percent.

Comparisons of predicted and measured response during pile load tests # 1, # 2, and

# 3 axe shown in Figures 6.14, 6.15, and 6.16, respectively. Changes in predicted

shear transfer during these tests are consistent with field measurements (Figures

6.14a, 6.14b, 6.15a, 6.15b, 6.16a and 6.16b). It should be noted that the shear transfer

is the average value along the shaft of the pile segment model. The estimations

of shear transfer at failure were successfully achieved. The predicted changes in

the effective radial stresses compare well with the measured changes (Figures 6.14c,

6.15c, 6.16c).

The estimated pore water pressure variations together with the field measure

ments zire presented in Figures (6.14d, 6.15d, 6.16d). Because of the difference be

tween the measured and predicted pore water pressures immediately after driving,

pore water pressures were predicted less than the measured values during tests #


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Stress path during pile driving (A-B)

Stress path during consolidation (B-C)

Yield surface

Phase change line20

10

00 100 200 300

J i , kPa

Figure 6.13: Predicted effective stress path for a soil element at the soil-pile interface during driving and soil consolidation around the 3 in. diameter pile segment model.

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5000 1000 1500 2000

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Figure 6.14: Comparison of predicted and measured response during pile load test #1 conducted on the 3 in. diameter pile segment model.

(figure con'd)


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161

1 and # 2. However, the simulation predicted the correct trend of the pore water

pressure change during these tests.

Pile load test # 4 was simulated at the completion of consolidation (degree of

consolidation, U=100%). Unlike in the previous tests, loading and unloading cycles

were simulated during this test. As presented in Figures 6.17a and 6.17b, both

the predicted and measured shear transfer show consistent patterns during loading-

unloading cycles. Figure 6.17c presents the comparison between measured and pre

dicted effective radial stresses during the test. The field measurements did not show

any significant variation during the cycles of the test. The simulated results and

the field measurements diverged at the end of the test. Changes in predicted and

measured pore water pressure during the test are shown in Figure 6.17d. The results

compared well with each other.

Finally, numerical simulation was performed to predict the behavior during the

final two-way cyclic pile load test. The test consisted of five tension-compression

cycles. The numerical results are in good agreement with the measurements, as

presented in Figure 6.18. The predicted shear transfer change with displacement,

in the first tension cycle, was slightly lower than in the field measurements. As

depicted in Figure 6.18a, estimation of shear transfer during the compression part

of the cycle is consistent with field measurements. During the tension part of the

cycle, the predicted shear transfer was slightly lower than the field measurements.

However, the numerical simulation accurately captured the response patterns during

the tension-compression cycles (Figure 6.18b).

Predicted and measured variations of the effective radial stresses during the test

are compared in Figure 6.18c. Field measurements did not show a consistent behavior

during the test. The predicted results were slightly lower than the measurements.

A comparison of the predicted and measured pore water pressure during the test is


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-2.0E-3 -1.0E-3 0.0E+0 1.0E-3 2.0E-3

Displacement, m


Figure 6.18: Comparison of predicted and measured response during the cyclic pile load test conducted on the 3 in. diameter pile segment model.

(figure con'd)

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167

depicted in Figure 6.18d. The variations of predicted pore water pressure during the

test cycles were slightly higher than field measurements.

Regardless of the slight differences between the measured and predicted responses,

the finite element simulation successfully estimated the various behavior patterns of

the field experiments.

6.6 Pile Setup

Conducting pile load tests at different time intervals during consolidation provided

the necessary means for finding the variation in pile shaft capacity with time. For

each conducted load test (field measurements and finite element simulation), the peak

and residual values of the shear transfer were identified. Figure 6.19 compares the

predicted and measured increases with time in the peak and residual shear transfer

during the pile load tests. Within the same set of measurements and predictions, the

differences between the peak and residual values were small.

The predicted peak shear transfer increased from 11 kPa, 54 minutes after driving,

to 19 kPa, at the end of consolidation. The corresponding field measurements were

increased from 10 to 24 kPa. Therefore, the increase in the shear transfer with time

occurred in both the measured and predicted results. In conclusion, even though the

simulation estimated a slightly lower setup at the end of consolidation, the predictions

compared very well with the field measurements, and both results have consistent

patterns.


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Field test, peak

Field test, residual30

— A -— Prediction, peak(0CL

— - Prediction, residual

u>c

I—(0a>xzCO

1.0E+3 1.0E+4 1.0E+5 1.0E+6

Time after pile driving, sec

Figure 6.19: Comparison o f predicted and measured increase in soii-pile shear transfer (setup) during pile load tests performed at different time periods after driving.

CHAPTER 7 NUMERICAL SIMULATION OF THE BEHAVIOR OF OPEN-ENDED PILES

7.1 Introduction

Open-ended piles are usually steel pipes or concrete cylinders. These piles are com

monly used in geotechnical and offshore engineering applications. It is essential to

study the behavior of open-ended piles in order to evaluate their response during

driving and loading.

As presented in Chapter 5 and 6 , the numerical simulation of the behavior of

closed-ended piles was successfully achieved. But the behavior of open-ended piles,

during driving and loading, differs from that of closed-ended piles. As the open-

ended pile penetrates the ground, the soil enters the pile. The accumulation of soil

inside the pile may result in pile plugging, which affects the behavior of the pile.

In this study, the method used for predicting the response of closed-ended piles

was modified for open-ended piles. This method was applied to predict the response

of field experiments (Earth Technology Corp. 1986) conducted on a 3 in. diameter

pile segment model with changeable cutting shoes, during driving, consolidation,

and load testing. The purpose of using the cutting shoes was to investigate the

behavior of open-ended piles. The numerical results were compared with the field

measurements in order to demonstrate the validity of the procedure for open-ended

piles.

This Chapter presents the results of the numerical simulation of the field exper

iments conducted on the 3 in. x 0.125 in. and 3 in. x 0.065 in. open-ended pile

segment models driven into Sabine clay.

169


170

7.2 Field Experim ents

Part of the Earth Technology Corp. study (1986) was to investigate the response of

open-ended piles during driving, subsequent consolidation, and load testing. There

fore, field experiments were conducted on a 3 in. diameter pile segment model driven

with cutting shoes of 0.125 in. and 0.065 in. wall thicknesses.

Changes in the measured total radial and pore water pressures as well as in the

corresponding calculated effective radial pressure, during soil consolidation for the 3

in. x 0.125 in. and 3 in. x 0.065 in. open-ended pile segment models, are shown in

Figures 7.1 and 7.2, respectively. Earth Technology Corp. (1986) reported a leakage

of moisture during the experiment on the 3 in. x 0.125 in. model, as a result,

instability in the recorded data was observed. The incident occurred at a degree

of consolidation of 93%; thus, the loss of the pressure data did not invalidate the

experiment. In the case of the 3 in. x 0.065 in. model, the calculated effective radial

pressure was negative during the first seventy minutes of soil consolidation. Earth

Technology Corp. (1986) stated that these values should not be considered real, but

rather indicative of nearly equal total and pore water pressures.

Pile load tests were carried out on the 3 in. x 0.125 in. model at 16%, 36%,

57%, 86%, and 95% degrees of consolidation, as described in Figure 7.3. For the

3 in. x 0.065 in. model, the pile load tests were conducted at different degrees of

consolidation, namely: 39%, 72%, 85%, and 99% (Figure 7.4). The measured shear

transfer changes with displacement for these tests are presented in Figures 7.5 and

7.6.

7.3 Finite Element M esh

The finite element mesh used for the simulation of consolidation and pile load tests

is shown in Figure 7.7. The 8-node axisymmetric solid element CAX8RP was used


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pre

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Pa

Tota

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dial

and

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, kP

a

171

400

200


Pore water pressure

10 1000 100001001

Time, min

(a) Total radial and pore pressures.

200

100

01000 1000010 1001

Time, min


Figure 7.1: Variation of measured pressures during soil consolidation around the 3 in. X 0.125 in. opend-ended pile segment model driven at Sabine Pass, Texas (after Earth Technology Corp., 1986).


Effe

ctiv

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dial

pre

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Pa

Tota

l ra

dial

and

po

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pres

sure

, kP

a

172

400

200

Pore water pressure


10 1001 1000 10000

Time, min

(a) Total radial and pore pressures.

100

50

0 -

1001 10 1000 10000Time, min


Figure 7.2: Variation of measured pressures during soil consolidation around the 3 in. X 0.065 in opend-ended pile segment model driven at Sabine Pass, Texas (after Earth Technology Corp., 1986).


173







Install 3 in. x 0.125 in. partial displacement pile segment model

Figure 7.3: Sequence o f the load tests conducted on the 3in. x 0.125 in. open-ended pile segment model.


174






Install 3 in. x 0.065 in. partial displacement pile segment model

Figure 7.4: Sequence o f the load tests conducted on the 3in. x 0.065 in. open-ended pile segment model.


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30

COCL■ * 20t _

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U = 16%

U = 36%

U = 57%

U = 86%

U = 95%-10

O.OE+O 1.0E-3 2.0E-3 3.0E-3 4.0E-3

Displacement, m

Figure 7.5: Results of tension pile load tests conducted on the 3 in. X 0.125 in. open-ended pile segment model (after Earth Technology Corp., 1986).

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30

U = 39%

U = 72%

U = 85%toCL

U = 99%

£(/>c£(00)x :(/)

-10 ----0.0E+0 1.0E-3 2.0E-3 3.0E-3 4.0E-3

Displacement, m

Figure 7.6: Results o f tension pile load tests conducted on the 3 in. X 0.065 in. open-ended pile segment model (after Earth Technology Corp., 1986).

ON

177

S>i

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< 3 x

<5 (N

Element 8-node ABAQUS-CAX8RP 782 elements 2473 nodes

Figure 7.7: Finite element mesh used to simulate soil behavior around the 3 in. X 0.125 in. and 3 in. X 0.065 in. pile segment models.


6

178

to form the mesh, which consisted of 782 elements and 2473 nodes. The soil mass

around the pile was modeled as the HiSS-^i material.

7.4 P ile Driving

Open-ended pile driving was simulated based on the strain path method (Chen and

Baligh, 1983). In the present study, a computer code was developed to predict

the strain field around open-ended piles during driving. Details of the numerical

procedure used in this code to simulate the driving of open-ended piles were described

in Section 4.3.3.

When an open-ended pile is driven into the ground, part of the surrounding soil

is displaced and another part enters the pile. The soil plug, or soil column inside

the pile, may affect the behavior of the pile during driving and loading. A pipe pile

is considered a closed-ended pile when sufficient internal shaft friction between the

soil plug and the pile is mobilized to provide end bearing over the full cross section

at the pile base (Leong and Randolph, 1991). The pile is then called fully plugged.

The totally unplugged pile is an open-ended pile, where the soil plug has no effect

on the behavior of the pile. Field measurements of the upper face of soil plugs inside

open-ended piles were conducted by Sovinc et al. (1985). Results indicated that the

piles remained mainly unplugged during driving. Leong and Randolph (1991) stated

that the assumption that piles remain essentially unplugged during pile driving is

generally acceptable.

Soil may plug open-ended piles during testing. Smith and Willson (1986) showed

that most piles plug during static loading. The soil plug may affect the behavior of

the pile differently under compression and tension loading. Beringen and Ruiter

(1987) stated that the plugged or unplugged condition of the pile is not significant

for pullout capacities. Most of the load tests simulated herein are static tension


179

tests. Therefore, open-ended pile driving and testing is simulated herein without

considering soil plug response.

7.4.1 T h e 3 in . x 0.125 in . O p en -en d ed P ile Segm ent M odel

Figure 7.8 shows the predicted effective stress path for a soil element at the soil-pile

interface during pile driving. The soil element is located at a depth of 16 m. This

element has an initial ff0-condition at point A before pile driving and a low stress

state at point B at the end of pile driving. Variations of the predicted effective stresses

with the normalized radial distance immediately after driving are shown in Figure

7.9. Examination of Figure 7.9 indicates that the predicted effective stresses were

significantly reduced due to pile driving. The reduction in these stresses occurred

around the pile up to a distance of about three times pile diameter. The estimated

effective radial stress at the soil-pile interface was substantially decreased from 69

kPa before, to 18 kPa after pile driving.

Based on field measurements, the Earth Technology Corp. (1986) calculated neg

ative effective radial stress after driving. It was suggested that these measurements

should not be considered real (Earth Technology Corp., 1986). The prediction of the

effective radial stress immediately after driving is higher than field measurements,

as shown in Figure 7.10.

The predicted pore water pressure distribution at the end of pile driving is shown

in Figure 7.11. The maximum predicted pore water pressure at the pile wall was

75% of the corresponding measured value. Therefore, the maximum excess pore

water pressure was underestimated by a factor of 0.5.

7.4.2 T h e 3 in . x 0.065 in . O p e n -e n d e d P ile Segm ent M odel

Numerical simulations of open-ended pile driving, subsequent consolidation, and

load testing are sim ila r for both the 3 in. x 0.125 in. and the 3 in. x 0.065 in.


180

30

a. 20JCaCM

10

00 100 200 300

Jn , kPa

Figure 7.8: Predicted effective stress path for a soil element at the soil-pile interface during driving and soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.


Yield surface

Phase change line

Effective stress path during pile driving

Effective stress path during consolidation

Effe

ctiv

e st

ress

, kP

a

181

150

125

zz

100

1 10 100


Figure 7.9: Predicted distribution of effective stresses at 16 m below the ground surface, immediately after driving 3 in. X 0.125 in. open-ended pile segment model.


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200

Field measurements

Finite element simulation(0£ 150tnco£co

100TJsa>>’■8I 50ud

After pile driving

1 10 100 1000 10000


Figure 7.10: Comparison of measured and predicted effective radial stresses during soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.

00N>

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400


Field measurementnsQ_- 300

23(ACO0)u .Q.

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1001 10 100


Figure 7.11: Variation with radial distance of predicted pore water pressure immediately after driving 3 in. X 0.125 in. open-ended pile segment model.

00

184

models. Therefore, the simulation results for the 3 in. x 0.125 in. model is discussed

in detail herein, while the results of the 3 in. x 0.065 in. are presented in Appendix

B of this report.


Finite element simulation of the consolidation phase started immediately after pile

driving. Initial conditions required for the simulation were obtained from the end of

driving stage. During soil consolidation, different pile load tests were simulated at

time intervals corresponding to the degrees of consolidation specified in Figures 7.3

and 7.4.

7.5.1 T he 3 in . x 0.125 in . O p en -en d ed P ile S eg m en t M odel

Changes in the predicted and measured effective radial stresses during the subsequent

consolidation are compared in Figure 7.10. Immediately after driving, the predicted

effective radial stress is higher than the field measurement. During consolidation,

the estimated variation of the predicted effective radial stress is slightly lower than

the field measurement. However, the results of the finite element simulation are

considered satisfactory.

During the consolidation phase, the excess pore water pressures gradually dis

sipated. The dissipation was associated with the increase in the effective stresses.

Predicted pore water pressure distributions at different time periods during con

solidation are shown in Figure 7.12. The corresponding estimated increase in the

effective radial stresses is presented in Figure 7.13. The predicted and measured

normalized excess pore water pressures are compared in Figure 7.14. The predicted

and measured dissipation patterns are consistent.


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180

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47409 sec 200000 sec

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100

Figure 7.12: Predicted dissipation of pore water pressures during soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.

OO

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1025 sec 2049 sec

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5889 sec

10497 sec 17409 sec

47409 sec 200000 sec

10 100

Normalized distance from pile wall (r/rD)

Figure 7.13: Predicted increase in effective radial stresses during soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.

OOON

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1.0E-1 1.0E+0 1.0E+1 1.0E+2 1.0E+3 1.0E+4


Figure 7.14: Comparison of predicted and measured normalized excess pore water pressures during soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.

00

188

When, the excess pore water pressures are fully dissipated, the maximum increase

in the effective stresses is predicted (Figure 7.15). In this study, as indicated by the

figure, the effective radial stress increased from 18 kPa immediately after driving to

96 kPa at the end of consolidation. Effective stresses close to the pile also increased

when compared to the initial Ka-conditions before pile driving. Predicted effective

radial stress near the pile increased from 69 kPa before pile driving, to 96 kPa at

the end of consolidation. On the other hand, the effective stresses at the far field

remained unchanged during consolidation.

The predicted effective stress path for a soil element at the soil-pile interface dur

ing consolidation is shown in Figure 7.8. The stress condition at point C represents


The static tension and final two-way cyclic pile load tests were simulated during

consolidation in accordance with the sequence described in Figure 7.3. The predicted

and measured behavior during pile load tests # 1, # 2, # 3, and # 4 are compaxed in

Figures 7.16, 7.17, 7.18, and 7.19. Variations of the predicted and measured shear

transfer during these tests are shown in Figures 7.16a, 7.16b, 7.17a, 7.17b, 7.18a,

7.18b, 7.19a and 7.19b. Comparisons of the predicted and measured shear transfer

variations indicate the presence of slight differences between these values at failure.

However, the results of the finite element simulation are in reasonable agreement

with field measurements.

The predicted and measured changes of effective radial stresses and pore water

pressures during these tests axe depicted in Figures 7.16c, 7.16d, 7.17c, 7.17d, 7.18c,

7.18d, 7.19c and 7.19d. The estimations axe slightly lower than the field measure

ments. Variations of predicted pore water pressures during these tests show dis

crepancies with the experimental results. The differences between the finite element


189

100

(0CL

(A<0(1)(0<D>o£ID zz

00

101 100Normalized distance from pile wall (r/ro)

Figure 7.15: Predicted distribution of effective stresses at a depth of 16 m at the end of soil consolidation around the 3 in. X 0.125 in. open-ended pile segment model.


Shea

r tr

ansf

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Shea

r tr

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190

20

10

0

Degree of consolidation. U = 16%

Reid pile load test

Rnite element simulation-10

0.0E+0 5.0E-4 1.0E-3 1.5E-3

Displacement, m


20

\__


Field pile load test-10


0 40 80 120Time, sec


Figure 7.16: Comparison of predicted and measured response for pile load test # 1 conducted on the 3 in. X 0.125 in. open-ended pile segment model.

(figure con'd)


Pore

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191

100




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Time, sec

(c) Effective radial stress vs tiine.

400

300

200 -

10040




80 120

Time, sec



Shea

r tr

ansf

er,

kPa

Shea

r tr

ansf

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kPa

192

20




-20 ----0.0E+0 1.5E-35.0E-4 1.0E-3

Displacement, m





-203002001000

Time, sec



(figure con'd)


Pore

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pre

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e, k

Pa

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str

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193

100

75 -

------------------------------------50 -

25 -Degree of consolidation, U = 36%



100 200

Time, se c

300


300

200 -

100


Reid pile load test


100 200

Time, sec

300



Shea

r tr

ansf

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Shea

r tr

ansf

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194

20

10

0




-20 ----O.OE+O 1.0E-35.0E-4 1.5E-3

Displacement, m


20

10

0




0 200 400 600 800Time, s e c



(figure con’d)


Pore

w

ater

pre

ssur

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Effe

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str

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195

100 -

50 -


Field pile loasd test


200 400

Time, se c

600 800


300

200 -

100200




I _ i_

400

Time, sec

600 800



Shea

r tr

ansf

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Shea

r tr

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196

30

20

10




-20 ----O.OE+O 1.0E-35.0E-4 1.5E-3

Displacement, m


30

20

10




0 150 300 450

Time, sec

(b) Shear transfer vs time


(figure con'd)


Pore

w

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pre

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Effe

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197

250

200

150

100

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00 150 300 450

Time, sec


250

200

150

100

0 150 300 450

Time, sec





i________I________■ ..




v___


198

solution and the field measurements decrease as the excess pore water pressures dis

sipate. However, these differences did not affect the predictions of the shear transfer

during the simulation of the pile load tests.

Pile load test # 5 was simulated at 95% degree of consolidation. The test con

sisted of both loading and unloading cycles. Figure 7.20a shows the predicted and

measured variations of the shear transfer with the displacement. The numerical re

sults are lower than the field measurements. However, the behavior is well predicted

during the unloading-reloading cycles. Variations of the predicted and measured

shear transfer during the test are comparable, as shown in Figure 7.20b.

Figure 7.21a presents the variations of the estimated effective radial stress during

the test. Because of the leakage that occurred near the end of the experiment, the

measured total radial and pore water pressures were not considered for comparisons.

Therefore, only the finite element results are presented. The change in predicted

pore water pressure with time is shown in Figure 7.21b.

The numerical simulation of the final two-way cyclic pile load test was performed

near the end of consolidation. The test consisted of four tension-compression cycles.

The finite element results are in good agreement with the field measurements, as

presented in Figures 7.22 and 7.23.

The estimated shear transfer variations during the cycles axe slightly lower than

the measured values. However, the predicted and measured results are consistent,

and the numerical simulation accurately captured the response patterns during the

tension-compression cycles.

The predicted variations of effective radial stress and pore water pressure during

the test are shown in Figures 7.23a and 7.23b. The field measurements are not

presented due to their fluctuations during the test.


199

40

(0CL 20

l—<Dcnc2L_(00)r.CO




1.0E-3 2.0E-3 3.0E-3O.OE+O

Displacement, m


CDQ.

inc(0k_

i—(0d)x:CO




-301000 2000 30000

Time, s e c




Pore

w

ater

pre

ssur

e, k

Pa

Effe

ctiv

e ra

dial

str

ess,

kPa

200

100

1000 2000o 3000

Time, se c


300

100 1 1 1 1 1------------0 1000 2000 3000

Time, se c


Figure 7.21: Predicted response during pile load test # 5 performed on the 3 in. X 0.125 in. open-ended pile segment model.


Shea

r tr

ansf

er,

kPa

201

40

20

0

-20



O.OOE+O 1.00E-3 2.00E-3-1.00E-3

Displacement, m


Figure 7.22: Comparison of predicted and measured response for the final cyclic pile load test conducted on the 3 in. X 0.125 in. open-ended pile segment model.

(figure con'd)


Shea

r tr

ansf

er,

kPa202

40

20

0

-20



0 1000 2000 3000 4000Time, sec



Pore

w

ater

pre

ssur

e, k

Pa

Effe

ctiv

e ra

dial

str

ess,

kPa

203

100

75

50

2520000 1000 3000 4000

Time, sec


250

200

150

1000 2000 3000 40000

Time, sec(b) Pore water pressure vs time.

Figure 7. 23: Predicted response for the final cyclic pile load test conducted on the 3 in. X 0.125 in. open-end pile segment model.


204

7.5.2 The 3 in . x 0.065 in. O pen-ended P ile Segm ent M odel

Finite element simulation of the subsequent consolidation and pile load tests was

conducted to predict the response of the 3 in. x 0.065 in. open-ended pile segment

model. The estimated behavior of the model was compared with the results of the

field experiments. In order to avoid repetition, comparisons of the numerical results

with measurements are presented in Appendix B.

7.6 P ile Setup

Examination of Figures 7.16, 7.17, 7.18, 7.19, 7.20 and 7.22 indicates that the finite

element results are generally consistent with the field measurements. The predicted

variation of shear transfer during these tests is shown in Figure 7.24. The magnitude

of the estimated shear transfer at failure increased with the increase in the degree

of consolidation. The finite element successfully predicted the increase in the shaft

capacity with time (the pile setup).

The predicted and measured shear transfer values at failure are compared in

Figure 7.25. The comparison includes the peak and residual shear transfer values.

Within the estimated and measured values, the difference between the peak and the

residual is small. Finite element solutions axe consistent with the measured increase

with time in the peak and residual shear transfer values. At the end of consolidation,

the predictions axe slightly lower than the field measurements.

The response of the open-ended pile during pile driving, subsequent consolida

tion, and pile load tests was predicted using numerical and finite element methods.

The numerical solutions for the different life stages of the open-ended pile axe suc

cessfully achieved. These solutions are found to be consistent with the field measure

ments.


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without

permission.

20

15

10U = 95%

U = 86%

U = 57%5

U = 36%

U = 16%

0 It—0.0E+0 5.0E-4 1.0E-3 1.5E-3 2.0E-3

Displacement, m

Figure 7.24: The finite element prediction of the increase in shear transfer for pile load tests conducted during soil consolidation around the 3 in. X 0.125 in. pile segment model.

toO

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(0CLyc.uT£.<nc2i .(0Q)x z(/)

50

Field test, peak

Field test, residual40

— A- — - Prediction, peak

— Q- — - Prediction, residual30

20

10

0 ---1.0E+1 1.0E+2 1.0E+3 1.0E+4


Figure 7.25: Comparison of predicted and measured pile setup for the 3in. X 0.125 in. open-ended pile segment model.

NJOO N

CHAPTER 8 NUMERICAL EXPERIMENTS ON A FULL-SCALE CLOSED-ENDED PILE

8.1 Introduction

Numerical simulations of the behavior of closed-ended as well as open-ended pile

segment models during driving, consolidation, and load testing were successfully

conducted. The numerical study was performed on pile segment models due to

the availability of the field experiments necessary for the verification and evaluation

schemes. However, the method is also applicable for investigating the behavior of

full-scale friction piles driven into saturated clay. For demonstration, finite element

numerical simulation of the behavior of a full scale pile was earned out.

This chapter presents the results of the numerical simulation of the various life

stages of a full-scale friction pile driven into saturated clay.

8.2 Pile Characteristics

A full-scale pile, 0.5 m in diameter and 10 m in length, was considered in order to

perform various numerical experiments. The material parameters, soil conditions,

and numerical stability and convergence parameters of Sabine clay were used in the

simulation.

The soil mass was discretized according to the finite element mesh presented

in Figure 8.1. The mesh consists of 377 elements and 1216 nodes. The 8-node

axisymmetric solid element CAX8RP was used to form the finite element mesh. The

soil mass around the pile was modeled as the HiSS- material.

207


k5 m 208

Element 8-node ABAQUS-CAX8RP 377 elements 1216 nodes

Figure 8.1: The finite element mesh used to discretize the soil around the pile.


209

8.3 Simulation o f Pile Driving

The pile was driven at the ground surface. As a result, the soil mass around the pile

was assumed to be affected by pile driving. The simple pile approach (Baligh, 1984),

developed for deep penetration of closed-ended piles, was utilized to simulate pile

driving.

At the end of driving, the effective stresses of the soil around the pile significantly

decreased. In contrast, the effective stresses of the soil in the far field remained under

AT,,-conditions. Figure 8.2 shows that low magnitudes of predicted effective stresses

around the pile extend to about ten times pile diameter. The effective radial stress

decreased from about 50 kPa before driving to 12 kPa immediately after driving.

The predicted pore water pressure distribution at the end of pile driving is de

picted in Figure 8.3. Immediately after pile driving, excess pore pressure developed

in the soil around the pile, up to a distance of about ten times pile diameter.


The initial conditions required for the finite element simulation of consolidation were

obtained at the end of driving. They include stresses and pore water pressures

that satisfy equilibrium, strain compatibility, and constitutive relations. The finite

element simulation was conducted based on the nonlinear transient analysis available

in ABAQUS.

The effective stresses gradually increased as the excess pore water pressures dis

sipated during consolidation. Dissipation of the excess pore water pressure during

soil consolidation around the pile is presented in Figure 8.4. Most of the excess pore

water pressures dissipated within 1000 hours after driving. Konard and Roy (1987)

reported that the excess pore water pressures generated during driving 22 cm di

ameter steel piles were dissipated 600 hours after driving. For the 0.5 diameter pile


Effe

ctiv

e st

ress

, kP

a

210

80

60

40

20

0

Figure 8.2: Predicted distribution of effective stresses immediately after driving 0.5 m diameter full-scale pile.

□ O O D D I

o o o c o

1 10

Normalized distance from pile wall (r/r)


Pore

w

ater

pre

ssur

e, k

Pa211

120

100

80

60101


Figure 8.3: Predicted distribution of pore water pressure immediately after driving 0.5 m diameter full-scale pile.


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140 1 1—i—i i n~ 1 ----------- 1-------- 1— i— i i i i

(0CL 120

S3 w tn 9>

<oCD$a>L -oQ.

100

80

60 1.0E+0

_i i i i i i 11 j i i i i i 111 j i i i 1111 j i i i i 111 1

1.0E+1 1.0E+2 1.0E+3

Time after pile driving, hours

1.0E+4

Figure 8.4: Variation with time of predicted pore water pressure during soil consolidation around the 0.5 m diameter full-scale pile.

212

213

simulated herein, it is therefore acceptable for most of the excess pore water pres

sures to dissipate during the 1000 hours. Furthermore, the pattern of the dissipation

curve is similar to those obtained in the field by the Earth Technology Corp. (1986).

The estimated increase in the effective radial stress during consolidation is de

picted in Figure 8.5. At the end of consolidation, the maximum increase in the

effective stresses developed with complete dissipation of the excess pore water pres

sures. The effective radial stress increased from 12 kPa immediately after driving to

53 kPa at the end of consolidation.

Four static compression load tests were simulated during the consolidation phase.

These tests were simulated at 10, 43, 78, and 100 percent degrees of consolidation.

For each pile load test, linear displacement was applied during the test until the

shear transfer at the soil-pile interface became constant with time. Variations of

shear transfer with displacement for these tests Me presented in Figure 8 .6 . The

predictions of shear transfer change with displacement axe consistent with those

obtained in the previous chapters.

8.5 P ile Setup

Simulating pile load tests at different time intervals during consolidation provided the

necessary means to estimate pile setup. During the pile load test, the maximum shear

stress at the soil-pile interface represents the maximum mobilized skin friction for

resisting pile loading. For these simulated tests, the variations with displacement of

the average skin friction along the pile shaft Me shown in Figure 8 .6 . The maximum

average value for each pile load test was identified (Figure 8 .6 ). Shown in Figure

8.7 is the variation with time of the average skin friction along the pile shaft for

these load tests. Examination of Figure 8.7 indicates that the bearing capacity of

the pile increased during soil consolidation and a pile setup developed. The average


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_i i i i i i 11 J i i i i i 111 i i i i i i n J I 1- 1 L l l l l

1.0E+0 1.0E+1 1.0E+2 1.0E+3

Time after pile driving, hours

1.0E+4

Figure 8.5; Variation with time of predicted effective radial stress during soil consolidation around the 0.5 m diameter full-scale pile.

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(0CL

i t fros zV)a>£o>co(0co

* 3oEc12<na>a)2a>><

10

5

Degree of consolidation

U = 100%

U = 78%

U = 43%0U = 10%

0E+0 2E-3 4E-3 6E-3

Displacement, m

Figure 8.6: Variations with displacement of average skin friction along the shaft for different pile load tests conducted during soil consolidation around the 0.5 m diameter full-scale pile.

NJ •—*

Ave

rage

sk

in fri

cition

alo

ng

the

shaf

t, kP

a

216

10

8

6

4 -----1.0E+0 1.0E+1 1.0E+2 1.0E+3 1.0E+4

Time after pile driving, hour

Figure 8.7: Predicted increase in average skin friction along the shaft of the pile during soil consolidation around the 0.5 diameter full-scale pile.


217

skin, friction increased from about 5 kPa six hours after driving to about 7.8 kPa at



CHAPTER 9 SUMMARY, CONCLUSIONS, AND

RECOMMENDATIONS

9.1 Summary

In this study, a general method for predicting pile setup, or the increase in bearing

capacity of friction piles driven into saturated clay, was developed based on consis

tent formulations during pile driving, consolidation, and loading. The behavior of

friction piles driven into saturated clay was investigated during pile driving, sub

sequent consolidation, and load testing. During soil consolidation, the response of

the pile upon loading provided the quantitative estimate of the variation of the pile

bearing capacity over time. The method was applied to both partial-displacement

(open-ended) and full-displacement (closed-ended) piles.

An HiSS constitutive interface model named was developed in order to capture

soil behavior under severe shear deformation at the soil-pile interface. The model

could also be used to characterize the behavior of the soil in the far field. The

developed model was implemented into the ‘strain to stress’ algorithm as well as into

the finite element program ABAQUS, in which the theory of nonlinear porous media

is used to perform the analysis.

The strain path method was used to simulate pile driving. The simple pile ap

proach was utilized for full-displacement (closed-ended) piles, while the concept of

the ‘ideal’ open-ended pile was utilized for partial-displacement (open-ended) piles.

A computer code was developed to obtain the strain field in the surrounding soil re

sulting from pile driving. This code used the proposed interface model to predict

the stress field in the soil around the pile during driving. Numerical simulations of

pile driving were performed in order to predict the field behavior of well-documented

218


219

field experiments conducted on pile segment models at Sabine Pass in Texas. Simu

lations included closed-ended as well as open-ended piles.

The finite element simulation of soil consolidation around the pile was conducted.

During the consolidation phase, pile load tests were simulated at different time inter

vals. The purpose of simulating these tests was to predict the variation of pile bear

ing capacity with time. These simulations were conducted using full-displacement

(closed-ended) as well as partial-displacement (open-ended) pile segment models.

The results of the finite element analysis were evaluated by comparing them with

pile behavior measurements in the field.

In this study, numerical simulations were performed on pile segment models due

to the availability of the field experimental tests necessary for the verification and

evaluation schemes. However, the method was also used to investigate the behavior

of full-scale friction piles driven into saturated clay. Finite element numerical exper

iments were conducted using a full-scale friction pile driven into saturated clay. The

numerical simulation included the various life stages of a full-scale driven friction

pile (pile driving, subsequent consolidation, and load testing). The variation of pile

bearing capacity with time was also investigated.

9.2 Conclusions

The following general conclusions are derived from this study:

• The ability of the general method developed for predicting the pile setup of

friction piles driven into saturated clay was demonstrated. The finite element

results were verified with respect to field experiments conducted on pile segment

models.

• Successful predictions of pile setup for both partial-displacement and full-

displacement pile segment models were achieved.


220

• Numerical experiments showed that the method is applicable for the prediction

of the pile setup of full-scale piles.

• Investigating the behavior of the pile during its various life stages resulted in

the reasonable evaluation of the pile bearing capacity.

• Finite element simulations of different pile load tests during soil consolidation

provided quantitative measures for the variation of pile bearing capacity over

time.

• An HiSS interface model named model was developed and evaluated with

respect to field experiments. This model was successfully used to predict soil

behavior under severe shearing at the soil-pile interface.

9.3 Recom m endations

The following are some recommendations for future research:

• Incorporating the effect of thixotropy on the pile setup.

• Further research to modfiy the constitutive model and to consider the changes

in the soil fabric during a loading (consideration of the micromechanical be

havior in the constitutive model).

• Studying the pile setup for a pile group.

• Investigating the effect of soil-plug response on the behavior of open-ended

piles and on the setup.

• Pefroming numerical experiments on full-scale piles with different dimensions

and considering different soil conditions. This can initiate a pile setup design

charts.


221

• Conducting an experimental program to obtain stresses and pore water pres

sures in the soil during pile driving, subsequent consolidation, and loading.

These measurements will provide the necessary means to evaluate the predic

tions of the stresses and pore water pressures away from the pile.


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2 27

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APPENDIX A PILE SETUP RELATED STUDIES

234


2 3 5

Table A .l: Pile setup related case studies.

Project Location and1 Reference(

San Francisco Bay, San Francisco, California,Seed and Reese (19SS).

Drammen, Norway,Eide et al. (1961).

Piles Used in the Project Full scale instrumented pipe piles, L=20 to 22 ft , D=6.0 in.

Full scale timber pile, D=0.35 m, L=13.1 m.

Loading Test Type Static pile load tests. Static short-term and long-term pile load tests .

Field Measurements Subsurface exploration (boreholes), vane shear tests.

Subsurface exploration, vane shear test, CPT, pore water pressure measurements with piezometers.

Laboratory Tests Physical properties (LL, PL, PL w, unit weight), unconfined compression test, one dimensional consolidation te st

Physical properties (LL, PL, PL w, unit weight, grain size distribution), salt content, unconfined compression test, standard oedometer consolidation te s t

Soil Conditions Bay mud (dark organic silty clay containing shells), nonsensitive soil.

Normally consolidated silty marine clay (sensitive clay where sensitivity ranges from 4 to 8).

Pile Setup Related Studies Measurements of the pile capacity at different times after driving, measurements of the shear strength and water content of the soil adjacent to the pile at different times after installation.

Performing pile load tests at different time intervals after driving. Determination of the final bearing capacity of the pile based on theoretical classical methods.

(table con’d)


2 3 6

Project Location and Reference

Nitsund Bridge, Norway,Skrede (1967), and Flaate (1971) and (1968).

Lower A rrow Lake, British Columbia, Canada,McCammon and Golder (1970).

Piles Used in the Project Full Scale Timber Piles,Dl=0.37 m, Ll=11.7 m, D2=0J3 m, L2=13.7 m.

Full scale open-ended steel pipe piles, D=24in, L=153ft (this pile was driven 5 more feet as closed-ended pile).

Loading Test Type Static pile load tests . Static pile load tests.

Field M easurements Swedish fall cone test, pocket vane shear test

Subsurface exploration (boreholes), field vane shear test

Laboratory Tests Physical properties (LL, PL, PI, w, unit weight), unconfined compression te s t .

Physical properties (LL, PL, PI, w), unconfined compression test

Soil Conditions Silty clay with thin silt layers. Stiff to very stiff grey silty clay.

Pile Setup Related Studies Performing pile load tests at various time intervals after driving.

Measurements o f the pile shaft capacity at different time periods after driving.

(table con’d)


2 3 7


Study a t Nova Scotia Technical College in Halifax,Canada,Clark and Meyerhof (1972).

Sabine Pass, Texas,Bogard and Matlock (1979),Earth Technology Corp. (1986), Wathugala (1°90), Katti 1990 Wathugala and Desai (1993), and Desai et al. (1993).

Piles Used in the Project

Steel pipe pile model, D=3in, t=0.1 in, L=30in.

Instrumented model pile segments with exchangeable cutting shoes, Dl=3.0in, D2=1.72in.

Loading Test Type Static pile load tests with different procedures.

Static and cyclic pile load tests.

Field Measurements Tests are performed in instrumented clay bed, measurements o f linear displacement, pore water pressure, and total stresses, vane shear tests.

Subsurface exploration, monitoring of the total lateral stresses, pore water pressure, shear transfer and relau've pile load soil displacement during driving, consolidation and load tests. Soil testing with self boring pressuremeter. Undrained shear strength by means o f torevane and miniature vane tests.

Laboratory Tests Physical properties (LL, PL, PI, grain size distribution, specific gravity), unconfined compression test, consolidated undrained triaxial compression test, drained direct shear test

Physical properties (LL, PL, PI, w, unit weight), mineralogical composition,UU triaxial tests, CU triaxial compression and extension tests with pore water pressure measurements, hydrostatic compression tests, one dimensional consolidation te s t

Soil Conditions Clay bed composed of illite with quartz and feldspar.

Highly plastic normally consolidated clay (sensitivity= 2.0 to 3.0).

Pile Setup Related Studies

measurements o f the undrained shear strength (unconfined compression tests, and vane shear tests) o f the soil adjacent to the pile shaft at different distances and time intervals after driving. Monitoring time variation of pore water pressures in soil near to the pile shaft at various distances and depths, measurement o f the total stresses in the soil in the vicinity of the pile, pile load tests on the model pile at different time intervals after driving.

Measurements o f radial stresses, pore water pressures, shear transfer and pile displacement during driving, consolidation and loading tests o f the piles. Finite element dynamic analysis o f nonlinear porous media with soil modeled as HiSS material are utilized to predict the shear transfer, stress state around the pile, pore water pressures and load tests.

(table con’d)


2 3 8


Empire, Louisiana,Cox and Kraft (1979), Kraft et al. (1981), Azzouz and Baligh (1984), Azzouz (1986), Bogard and Matlock (1990), and Azzouz et al. (1990).

Thorbum and Rigden (1980).

Piles Used in the Project Instrumented pile model, Dl=3.0in , D2=1.72in, Piezo-Laterai Stress Cell, full scale steel pipe piles, D=14.0in.

Full scale precast solid concrete square piles, LI =32 m, L2=29 m, B=0.25 m. circular pile, L=I5.24 m, D=525 m.

Loading Test Type Static pile load tests, cyclic axial load tests .

Static constant rate penetration pile load tests.

Field Measurements Subsurface exploration (boreholes), cone resistance and pore water pressure measurements by piezocone penetrometer, measurements of pore water pressure, lateral total stress during and after installation by means of the Piezo-Lateral Stress Cell, measurements o f lateral stresses, pore pressures, and shear transfer using the instrumented pile model.

Field vane shear strength test, CPT.

Laboratory Tests Index classification tests (PL, LL, PI, w, specific gravity), salt content, pH, organic matter, mineralogical composition, standard one dimensional consolidation test, one dimensional consolidation test with the MIT lateral stress oedometer cell, constant head pe rm eab ility test, consolidated undrained triaxial compression and extension tests with pore water pressure measurements, direct simple shear tests.

Physical properties, consolidated undrained triaxial compression test.

Soil Conditions Uniform firm grey clay, stiff to very stiff grey clay, (overconsolidated clay).

Slightly overconsolidated soft alluvium (considered as NC clay).

Pile Setup Related Studies Measurements o f the time variations of pore water pressures, lateral stresses on the pile shaft during and after installation by the Piezo-Lateral Stress Cell, prediction o f pile capacities based on field measurements and theoretical investigations, measurements o f radial stresses, pore water pressures, shear transfer and pile displacement during driving, consolidation and loading tests of the pile, development o f correlation for time dependent shear transfer.

Field measurements of the pile bearing capacity at different time intervals after driving, determination o f the shaft capacity using various techniques.

(table con’d)


239


Saint Alban, Quebec, Canada,Roy et al. (1981) and (1982), Roy and Lemieux (1986), and Konard and Roy (1986).

Jones Island Project, Milwaukee, Wisconsin,Fellenius et al. (1988), and Riker et al. (1988).

Piles Used in the Project Instrumented closed ended steel pipe piles, D=0.22 m, t=.08 m, L=7.6 m.

Steel Piles: thin wall pipes 12.75x0.375 in, H-piles 12HP63, heavy wall pipes, Iength= 110 to 156 ft.

Loading Test Type Static pile load tests. Dynamic monitoring with Pile Driving Analyzer during initial writing and restriking, static pile load tests.

Field M easurements Subsurface exploration (boreholes), vane field test, measurements o f pore water pressures during driving by means o f pore pressure cells, piezometers and piezocone penetrometer, measurements of total lateral stress during driving.

Subsurface exploration (boreholes).

L aboratory Tests Physical properties (LL, PL, PI, w, clay content), salt content, unconsolidated undrained triaxial compression test

Classification tests, unit weight, triaxial compression tests.

Soil Conditions Over consolidated soft silty clay o f marine origin underlain by soft clayey silt (sensitive clay).

Earth fill underlain by soft to medium stiff compressible postglacial silty clay and clayey silt with organics.

Pile Setup Related Studies Measurements o f the shaft bearing capacities o f tested piles at different times after installation from load tests, monitoring o f the pore water pressure dissipation with time after installation. Evaluation of the ultimate shaft bearing capacity based on empirical and theoretical methods.

Determination of piles capacities by means o f CAPWAP analysis and static pile load tests at different times after piles driving.

(table con’d)


2 4 0


1-310, New Orleans, Louisiana,McManis et al. (1988).

Alborg, Denmark,Skove and Denver (1988).

Piles Used in the Project Prestressed concrete cylindrical piles, Ll=84ft, Dl=54in, tl=5in, L2=84ft, D2=36in, t2=5in. prestressed concrete square Piles, Ll=84, Bl=30in, L2=84, B2=24in.

Prefabricated concrete square piles. Ll=19m, L2=21m, B=0.25m.

Loading Test Type Dynamic measurements with the Pile Driving Analyzer, static quick load tests, static pile load tests.

Static pile load tests, dynamic stress wave measurements.

Field Measurements Subsurface exploration (boreholes).

Subsurface exploration (boreholes).

Laboratory Tests Physical properties(LL, PL, PI, w), unconfined compression te s t .

Soil Conditions Soft to stiff grey clay. Sand layer underlain by thick clay layer.

Pile Setup Related Studies Conducting a series o f static pile load tests over different time intervals after driving to determine the change in the shaft capacity with time, determination o f the time variation o f the bearing capacity from the Pile Priving Analyzer measurements.

Measurements o f the pile bearing capacity at different time periods after driving based on static pile load tests and dynamic restrike o f the piles, CAPWAP analysis conducted for restrike blows to determine the shaft capacity from the dynamic stress-wave measurements.

(table con’d)


241


Harvey, LouisianaBogard and Matlock (1990)

Susquehanna River, New York. Chung (1988)

Piles Used in the Project Instrumented pile models, Dl=3.0 in, D2=1.72 in.

Full scale H-steel piles, (14HP73), L=30.5 m.

Loading Test Type Static pile load tests, cyclic axial pile load tests.

Field M easurements Record o f the total lateral stresses, pore water pressure, shear transfer and relative pile load soil displacement, cone penetration tests, to revane and miniature vane tests.

Subsurface exploration (boreholes), insitu vane shear test (undisturbed), insitu vane shear test (remolded).

Laboratory Tests Physical properties (LL, PL, PI, w), Unconsolidated undrained triaxial compression test, onedimensional consolidation test

Physical properties(LL, PL, PL w), Unconfined compression te s t , triaxial unconsolidated undrained tests.

Soil Conditions Soft to stiff grey normally consolidated clay with low plasticity.

Silty sand underlain by a thick layer of very soft to medium stiff normally consolidated clayey silt and silty clay, (sensitivity ranges between 3 and 12).

Pile Setup Related Studies Measurements of Radial stresses, pore water pressures, shear transfer and pile displacement during driving, consolidation and loading tests o f the pile. Development of correlation for time dependent shear transfer, comparisons of the rate of increase of the shear transfer with field measurements at two locations.

Monitoring the buildup and dissipation o f the excess pore water pressures after pile installation, evaluation of the excess pore water pressures based on methods proposed by randolph (1979) and Baligh (198S).


APPENDIX B NUMERICAL SIMULATION OF THE

BEHAVIOR OF 3 in. x 0.065 in. OPEN-ENDED PILE SEGMENT MODEL

242


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150

Field measurementscaO.■*. 100 10 <A

£(A

™ 50•os(U>O Q£LU


-501 10 100 1000 10000

Time after driving, minI

Figure B. 1: Comparison of measured and predicted effective radial stress during soil consolidation around the 3 in. X 0.063 in. open-ended pile sigment model.

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permission.

S>3inin

<0+ -»

s>oamin88TJa>Nro§o

Field measurements1.00 Finite element simulation

0.75

0.50

0.25

0.0010 100 1000


10000

Figure B.2: Comparison of predicted and measured normalized pore water pressures during soil consolidation around the 3 in. X 0.065 in. open-ended pile segment model.

244

Effe

ctiv

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ress

, kP

a

245

150

125

100

G— 0 Q QOOOOOroW OM nifflBBfr

10 1001


Figure B.3: Predicted effective stress distribution at 16 m below ground surface at the end of driving the 3 in. X 0.065 in. open-ended pile segment model.


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ission of the

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ner. Further


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toQ.JC2!3in<Aa)u .a&a?2oQ.

350


Field measurement300

250

200

1501 10 100


Figure B.4: Comparison o f predicted pore water pressure distribution and the maximum measured value immediately after driving the 3 in. X 0.063 in open-ended pile segment model.

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permission.

100

80

(0a.tnV)S(AraTJ2(U>

'■s

£tu

60

40

20

0

65 sec

1025 sec

2049 sec

3585 sec

5889 sec

10497 sec 17409 sec

47409 sec 200000 sec

10 100


Figure B.5: Variation of predicted effective radial stress with normalized distance from pile wall at different time periods during soil consolidation (at a depth of 16 m) around 3 in. X 0.065 in. open-ended pile segment model.

247

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260

65 sec240 1025 sec

2049 seccoCL

3585 sec2>3COco<DW

2205889 sec

10497 sec 17409 sec

a.200

£(0£ 47409 sec

200000 secO 180 Q.

160

1 10 100


Figure B.6: Variation o f predicted pore water pressure with the normalized distance from the pile wall at different time periods during soil consolidation around the 3 in. X 0.065 in. open-ended model pile segment.

400

Reproduced

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ission of the

copyright ow

ner. Further


without

permission.

(OCL

i _(00)•CCO Degree of consolidation, U = 39%



-10 —

O.OE+O 1.0E-3 2.0E-3 3.0E-3 4.0E-3

Displacement, m

Figure B.7: Comparison of predicted and measured shear transfer vs displacement for pile load test # 1 conducted on the 3 in. X 0.06S in. open-ended pile segment model.

249

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ission of the

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permission.

I

(0c2t _(00)JZU) Degree of consolidation, U = 72%



-10 ----0.0E+0 4.0E-4 8.0E-4 1.2E-3

Displacement, mI

Figure B.8: Comparison of predicted and measured shear transfer vs. displacement for pile load tests # 2 conducted on the 3 in. X 0.063 in. open-ended pile segment model.

toUio

Reproduced

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ission of the

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ner. Further


without

permission.

30

(0£L

t_r

<nc£L_CD0)•CCO Degree of Consolidation, U = 85%



-10 —

0.0E+0 5.0E-4 1.0E-3 1.5E-3 2.0E-3

Displacement, m

Figure B.9: Comparison of predicted and measured shear transfer vs. displacement for pile load test #3 conducted on the 3 in. X 0.065 in. open-ended pile segment model.

Reproduced

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ission of the

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ner. Further


without

permission.



Finite element simulation40

(0CLj*

30Wc

l _ro0)_ c

0)

0.0E+0 1.0E-3 2.0E-3 3.0E-3

Displacement, m

Figure B. 10: Comparison of predicted and measured shear transfer vs. displacement for pile load test #4 conducted on the 3 in. X 0.06S in. open-ended pile segment model.

252

VITA

Hani Hasan Titi was bom in Alarioub, Jordan on September 16, 1964. He com

pleted his bachelor of science degree in civil engineering from Yarmouk University,

Jordan, in June 1987. Then he attended Jordan University of Science and Technology

(JUST) from 1987 to 1989. He received his master of science degree in geotechnical

engineering in February 1990. During his study, he worked as teaching and research

assistant at the civil engineering department, JUST. He worked in the geotechnical

profession thereafter. He came to Louisiana State University in the fall semester of

1992. He is married to Eman Elayan. As of the date of completion this degree, he

has a beautiful girl, Miriam (10 months of age).

253


DOCTORAL EXAMINATION AND DISSERTATION REPORT

Candidate: Hani Hasan Titi

Major Field: civii Engineering

Title of Dissertation: THE INCREASE IN SHAFT CAPACITY WITHTIME FOR FRICTION PILES DRIVEN INTO SATURATED CLAY

Approved:

Major Professor and chairman

of -the Graduate SchoolDe<

EXAMINING COMMITTEE:

<2X v w \A A A

Date of Examination:

July 31, 1 996


the increase in shaft capacity with time for friction

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