b20592917.pdf - PolyU Electronic Theses

Abstract of thesis entitled

LABORATORY PULL-OUT TESTING STUDY ON SOIL

NAILS IN COMPACTED COMPLETELY DECOMPOSED GRANITE FILL

Submitted by

Su Li-Jun

for the degree of Doctor of Philosophy

at The Hong Kong Polytechnic University

in March 2006

Soil nailing is a technique for stabilizing soil slopes and excavations by installing a

large number of closely spaced passive inclusions into the in-situ soil mass. The soil

nailing technique has been increasingly used worldwide since its origination in the early

1970’s because of its technical and economical advantages. In Hong Kong, soil nailing

has been commonly used to stabilize new cut and sub-standard existing slopes since the

late 1980’s. The interface shear strength between a soil nail and the surrounding soil is a

key parameter for design and stability assessment of the soil nailing system. However,

in current practice in Hong Kong, this parameter is generally assumed to be the same as

the shear strength of the soil and verified by field pull-out tests in the construction stage.

Field verification tests are normally subjected to variations of the site conditions and the

results are therefore scattered. Laboratory pull-out tests have been carried out to help

overcome these problems and precisely investigate the factors influencing the nail-soil

interface shear strength. However, there were still some deficiencies in these tests and

can be improved.

A laboratory study of the pull-out shear resistance of cement grouted soil nails was

therefore conducted in compacted completely decomposed granite (CDG) fill. A

pull-out box with the internal dimensions of 1.0m in length, 0.6m in width and 0.83m in

height was designed and constructed to carry out the pull-out tests. An extension

cylindrical chamber was provided to house an extension part of the nail and ensure that

a constant 1.0m length of the test soil nail was maintained within the test box during

pull-out and no cavity would be left behind the end of the test nail. A waterproof front

cap was used to cover the soil nail head and prevent water leakage which made it

possible to apply back pressure to saturate the testing soil in submerged tests.

Comprehensive instrumentation was used and the earth pressure, suction, and pore

water pressure in the soil, the deformation of the testing soil, and the pull-out force and

displacement were measured. During the pull-out tests, the overburden pressure was

applied before drilling to simulate the actual construction procedure of the soil nailing

system.

A series of pull-out tests have been conducted using two copies of the above introduced

pull-out box. The test results showed that soil stresses around the hole were largely

released after drilling and recovery of the stresses due to grouting of the soil nail was

minimal. The development of pull-out shear resistance was mainly derived from the

constrained dilatancy of the soil. Tests in soil at different degrees of saturation showed

that the peak pull-out shear resistance varies with different degrees of saturation of the

soil, with higher resistances at the degrees of saturation of 50% and 75%. Pressure

grouting tests were carried out and showed that the average peak pull-out shear

resistance of the soil nail increased almost linearly with the increase in grouting

pressure. Numerical modeling was performed and agreements between the measured

and simulated results were good.

Acknowledgements

I wish to express my deepest gratitude to my chief supervisor, Professor J-H. Yin, for

his encouragement, support and guidance during this period of study. It was his endless

efforts and experienced guidance that made this work possible. The privilege of working

with Professor Yin has appreciably influenced my professional development and

perspectives.

Some of the tests in the study received financial support from Civil Engineering and

Development Department of The Hong Kong Special Administrative Region

Government and is gratefully acknowledged. The author would like to express thanks to

the Director of Civil Engineering and Development and the Head of the Geotechnical

Engineering Office for the permission of the use of data from those tests which received

financial support.

The improvement, setup and usage of the equipment and apparatus for the soil nail

pull-out resistance studies have received valuable comments from Mr. C. F. Chan, Mr. Y.

K. Shiu, Dr. S. L. Chiu, Dr. W. M. Cheung, Mr. W.K. Pun, Mr. Tony Cheung, Miss

Carrie Leung, Mr. K. L. Tang, and Mr. Danny Fu. All these comments are gratefully

acknowledged. I also wish to thank Mr. L. M. Chu, Miss W. H. Zhou and all the

technicians in the Soil Mechanics Laboratory of Department of Civil and Structural

Engineering in The Hong Kong Polytechnic University for their assistance in the setup

of the test apparatus and participation in some of the soil nail pull-out tests.

The author wishes to express his sincere gratitude to the two examiners, Professor R. J.

Jardine and Dr. L. M. Zhang for their invaluable comments in their thesis examination

reports and insightful questions and valuable suggestions during the oral examination.

I would like to express my special thanks and admirations to my wife, Xiao Jia, for her

understanding and support. I sincerely appreciate my parents and my sister for their

endless encouragement and constant support.

TABLE OF CONTENTS

CERTIFICATE OF ORIGINALITY

ABSTRACT

ACKNOWLEDGEMENTS

TABLE OF CONTENTS

LIST OF TABLES

LIST OF FIGURES

Chapter 1: INTRODUCTION

1.1 Background 1

1.2 Objectives 5

1.3 Organization of the thesis 5

Chapter 2: LITERATURE REVIEW

2.1 The soil nailing technique 10

2.1.1 Characteristics of soil nailing 12

2.1.2 Advantages and limitations of soil nailing 14

2.1.3 Fields of application 15

2.1.4 Soils suitable for soil nailing 17

2.2 Behaviour of soil nailing 17

2.2.1 Soil nailing mechanism 17

2.2.2 Nail-soil interface shear resistance 19

2.2.3 Influence of bending stiffness of the nail 22

2.2.4 Failure modes of soil nailed structures 23

2.3 Design methods for soil nailing structures 24

2.3.1 The Davis method 25

2.3.2 The French method 26

2.3.3 The German method 27

2.3.4 The Juran method 28

2.3.5 Discussion on current design guides and codes 30

2.4 Factors influencing the pull-out resistance 33

2.4.1 Soil conditions 33

2.4.2 Stress conditions 34

2.4.3 Methods of installation 34

2.4.4 The nail surface conditions 35

2.5 Research and development 35

2.5.1 Large scale model tests and field monitoring 35

2.5.2 Laboratory testing studies 38

2.5.2.1 Laboratory pull-out tests 38

2.5.2.2 Direct shear and interface shear testing studies 41

2.5.2.3 Centrifuge Modeling 42

2.5.2.4 Small scale tests 43

2.5.3 Numerical modeling 44

Chapter 3: EQUIPMENT AND APPARATUS FOR PULL-OUT TESTS

3.1 Problems studied by laboratory pull-out tests 60

3.2 Numerical study on boundary effect for design of the box 62

3.3 Design and construction of two pull-out boxes 64

3.3.1 Investigations to be conducted using the boxes 64

3.3.2 Description of the pull-out box 65

3.3.2.1 Extension cylindrical chamber covering the

soil nail end 67

3.3.2.2 A waterproof front cap to covering the soil nail head 68

3.3.2.3 Application of back pressure for saturation of the soil 69

3.4 Measures for reducing side friction of the box 70

3.5 Instrumentation and measurements 72

3.6 Drilling machine and cement grouting tools 75

3.6.1 Drilling machine 75

3.6.2 Equipment for cement grouting without and with pressure 76

3.7 Setup of the box for soil nail pull-out testing 79

3.8 Summary and conclusions 79

Chapter 4: MATERIAL PROPERTIES AND TEST PROCEDURES

4.1 Introduction 98

4.2 Material properties 98

4.2.1 Basic properties of the CDG soil 98

4.2.2 Determination of the shear strength of the soil 99

4.2.2.1 Conventional triaxial tests on saturated

soil specimens 99

4.2.2.2 Double cell triaxial tests on unsaturated

soil specimens 101

4.2.3 Properties of the cement grout 102

4.3 Calibration of transducers 103

4.4 Soil preparation and test procedures 105

4.4.1 Soil preparation 105

4.4.2 Preparation of soil specimens for triaxial tests 106

4.4.3 Application of vertical overburden pressure 106

4.4.4 Hole drilling and cement grouting 107

4.4.5 Installation of tensiometers and/or porewater

pressure transducers 107

4.4.6 Saturation of the test CDG soil 108

4.4.7 Pull-out of the nail 108

Chapter 5: INFLUENCE OF OVERBURDEN PRESSURE ON SOIL NAIL

PULL-OUT BEHAVIOUR AND RESISTANCE

5.1 Introduction 127

5.2 Stress variations during drilling and grouting 128

5.2.1 Stress release during drilling 128

5.2.2 Variation of earth pressure during and after grouting 130

5.3 Development of earth pressure during pull-out 131

5.4 Pull-out shear stress-displacement behaviour 133

5.5 Influence of overburden pressure on pull-out shear resistance 134

5.5.1 Peak pull-out shear resistance 134

5.5.2 Apparent coefficient of friction 135

5.6 Shear stress distribution on the nail-soil interface 136

5.7 Summary 139

Chapter 6: INFLUENCE OF DEGREE OF SATURATION ON SOIL NAIL



6.2 Earth pressure and pore pressure responses

during saturating the soil 153

6.3 Variations of earth pressure 155

6.3.1 Decreased earth pressure immediately after grouting 155

6.3.2 Variations of earth pressure 156

6.4 Failure patterns of the soil nail 157

6.4.1 Surface of the drillhole before and after pull-out 157

6.4.2 Failure surfaces of soil nails in the soil

at different degrees of saturation 157

6.5 Effect of degree of saturation of the soil on pull-out

behaviour and resistance 158

6.6 Summary and major findings 160

Chapter 7: EFFECT OF GROUTING PRESSURE ON SOIL NAIL PULL-OUT

BEHAVIOUR AND RESISTANCE


7.2 Variations of earth pressures 177

7.2.1 Variations of earth pressures during drilling

and pressure grouting 177

7.2.2 Variations of earth pressures during the whole

period of testing 179

7.3 Failure patterns of the soil nail 179

7.4 Influence of grouting pressure 180

7.5 Summary and conclusions 183

Chapter 8: NUMERICAL SIMULATION OF PULL-OUT TESTS


8.2 Simulation of the shearing plane 192

8.3 Description of the finite element model 193

8.3.1 Mesh and boundary conditions 193

8.3.2 Procedure of the simulation 194

8.3.3 Material properties 196

8.4 Simulation of the pull-out tests 198

8.4.1 Stress and strain rate contours 198

8.4.2 Variations of the vertical stress during the simulation 199

8.4.3 Influence of the overburden pressure 200

8.5 Parametric studies 201

8.5.1 Influence of dilation angle 201

8.5.2 Influence of grouting pressure 202

8.6 Summary 204

Chapter 9: SUMMARY, CONCLUSIONS AND SUGGESTIONS

9.1 Summary 219

9.2 Conclusions 221

9.3 Recommendations and suggestions 224

REFERENCES 226

LIST OF TABLES

Table 2.1 – Basic assumptions of different soil nailing design approaches 46

Table 4.1 – Properties of the CDG soil and cement grout 110

Table 4.2 – Shear strength parameters of the CDG soil 110

Table 8.1 – Material properties used in the finite element model 205

LIST OF FIGURES

Figure 1.1 – 1972 Sau Mau Ping Landslide 8

Figure 1.2 – 1972 Po Shan Road Landslide 9

Figure 2.1 – Equipment for launched soil nails (After Myles and Bridle1992)

47

Figure 2.2 – Comparison of soil nailing, micro piles and soil dowelling(After Bruce and Jewell 1986)

47

Figure 2.3 – Contrast of the construction sequence of reinforced earth and soil nailing (After Bruce and Jewell 1986)

48

Figure 2.4 – Soil nailing mechanism (After Byrne et al. 1998) 48

Figure 2.5 – Skin friction mobilization in pullout test (After Cartier andGigan 1983)

49

Figure 2.6 – Nails subject to shear and bending (After Mitchell 1987) 49

Figure 2.7 – Davis design method (After Shen et al. 1981b) 50

Figure 2.8 – Failure surfaces obtained by Davis design method and FiniteElement analysis (After Shen et al. 1981b)

50

Figure 2.9 – French design method (After Schlosser 1982) 51

Figure 2.10 – Final yielding curve for inclusion in French method (AfterSchlosser 1982)

51

Figure 2.11 – German design method: (a) Bilinear failure surface; (b)Acting forces and displacements; (c) Hodograph; and (d)Force polygon (After Gassler 1988)

52

Figure 2.12 – Juran’s kinematical limit analysis design method: (a)Mechanism of failure; (b) State of stress in inclusion; and (c)Theoretical solution for infinitely long bar adopted fordesign purpose (After Juran et al. 1990)

52

Figure 2.13 – Large model tests in the “Bodenvernagelung” project (AfterGassler 1992)

53

Figure 2.14 – Full scale test failure by breakage of inclusions in the“Clouterre” project (After Clouterre 1993)

53

Figure 2.15 – Full scale test failure by reducing adherence length of inclusions in the “Clouterre” project (After Clouterre 1993)

54

Figure 2.16 – Full scale test failure by excessive excavation in the“Clouterre” project (After Clouterre 1993)

54

Figure 2.17 – Pullout box used by Tei (After Tei 1993) 55

Figure 2.18 – Axial stress distributions along the nail obtained by Tei(After Tei 1993)

55

Figure 2.19 – Pullout box used by Pradhan (After Junaideen et al. 2004) 56

Figure 2.20 – Pullout box used by Chu (After Chu 2003) 56

Figure 2.21 – Relationship between effect of reinforcement and inclination of inclusion (After Jewell 1987)

57

Figure 2.22 – Failure surface of centrifuge model (After Tufenkjian and Vucetic 2003)

57

Figure 2.23 – Reduced scale model test of a nailed soil slope (AfterKitamura et al. 1988)

58

Figure 2.24 – Reduced scale model test of a nailed wall (After Kim et al. 1996)

58

Figure 2.25 – Mesh for (a) unreinforced slope and (b) soil nail slopedeveloped by Yang and Drumm (2000)

59

Figure 3.1 – A soil-nailed slope and pull-out box simulation 81

Figure 3.2 – Mesh for boundary effect study 81

Figure 3.3 – Vertical stresses induced by the hole drilling procedure 82

Figure 3.4 – Horizontal stresses induced by the hole drilling procedure 82

Figure 3.5 – Relationship between the vertical stress and the distance from the top surface of the drillhole in vertical direction for (a) elastic material and (b) Mohr-Coulomb material

83

Figure 3.6 – Relationship between the vertical stress and the distancefrom the top surface of the drillhole in vertical direction for (a) box size model (b) large size model

84

Figure 3.7 – Layout of transducers and pull-out equipment 85

Figure 3.8 – Design of the pull-out box – 3-D view 86

Figure 3.9 – Back and front views of the pull-out box 86

Figure 3.10 – Side view of the pull-out box 87

Figure 3.11 – Cross-section of the top cover and the pull-out box 87

Figure 3.12 – Watertight bolt 88

Figure 3.13 – Waterproof front cap 88

Figure 3.14 – Setup for saturating the soil 89

Figure 3.15 – Results of direct shear tests on the interface between the stainless steel sheet and the flexible plastic film withlubricating oil in between

89

Figure 3.16 – Method for reducing side friction 90

Figure 3.17 – Mesh and boundary conditions for investigating side friction of the box

90

Figure 3.18 – Vertical stress contour for small side friction 91

Figure 3.19 – Vertical stress variations with distance from the bottom of the model along the path in Figure 3.17

91

Figure 3.20 – Transducers and datalogger used in the tests 92

Figure 3.21 – Locations of the earth pressure cells 93

Figure 3.22 – Locations of the soil moisture probes (or pore pressuretransducers)

93

Figure 3.23 – Setup of the drilling machine 94

Figure 3.24 – Adjustment of drilling bit and drilling bars to ensure that they pass the centers of the two holes

94

Figure 3.25 – Drilling the hole 95

Figure 3.26 – Grouting without pressure (gravity head only) 95

Figure 3.27 – Grouting with pressure 96

Figure 3.28 – Setup of the box with full instrumentation and loading devices for soil nail pull-out under submerged condition

97

Figure 4.1 – Particle size distribution of the CDG soil 111

Figure 4.2 – Relationship between dry density and moisture content 111

Figure 4.3 – (a) Deviator stress vs. axial strain (b) pore water pressure vs. axial strain and (c) effective stress paths for conventionalCU triaxial tests with saturated soil specimens

112

Figure 4.4 – Relationship between s' and t for conventional CU triaxial tests with saturated soil specimens at axial strain of (a) 15% and (b) 20%

113

Figure 4.5 – (a) Deviator stress vs. axial strain (b) pore water pressure vs. axial strain and (c) effective stress paths for conventionalCD triaxial tests with saturated soil specimens

114

Figure 4.6 – Relationship between s' and t for conventional CD triaxial tests with saturated soil specimens at axial strain of (a) 15%and (b) 20%

115

Figure 4.7 – Schematic diagram of the double cell triaxial system 116

Figure 4.8 – The double cell triaxial system (After Yin 2003) 116

Figure 4.9 – (a) Deviator stress vs. axial strain (b) volume strain vs. axial strain and (c) s' vs. t for double cell triaxial tests on soil specimens at 38% degree of saturation

117


118


119

Figure 4.12 – Failed cement grout specimens of Uniaxial CompressiveStrength (UCS) tests

120

Figure 4.13 – Results of Uniaxial Compressive Strength (UCS) tests forthe cement grout (cylindrical specimen)

121

Figure 4.14 – Apparatus for calibrating earth pressure cells, pore water pressure transducers and transducers for the soil moisture probes

122

Figure 4.15 – Calibration results for the transducers 122

Figure 4.16 – Full bridge connection for strain gauge 123

Figure 4.17 – Soil compaction and pressure cell installation 123

Figure 4.18 – Taking soil samples for triaxial tests 124

Figure 4.19 – Tensiometer (Soil Moisture Probe) used in the test 125

Figure 4.20 – De-airing for a tensiometer 125

Figure 4.21 – A soil nail being pulled out under a dry soil condition 126

Figure 5.1 – (a) Total earth pressure and (b) vertical displacement vs. time during applying overburden pressure (OP) – for overburden pressure of 200kPa and initial degree of saturation (Sr) of 38%

140

Figure 5.2 – Total earth pressure vs. time during drilling and grouting –for overburden pressure of 200kPa and initial degrees of saturation (Sr) of (a) 38% and (b) 75%

141

Figure 5.3 – Average earth pressure and pull-out shear stress vs. (a) time and (b) pull-out displacement during pull-out – for overburden pressure of 300kPa and initial degree of saturation (Sr) of 75%

142

Figure 5.4 – Changes of average total earth pressure of P-Cells 1 to 4 at different stages of testing for tests with soil at 38% degree of saturation

143

Figure 5.5 – Relationship between average pull-out shear stress and pull-out displacement for tests in soil at degree of saturation (Sr) of 38%

143


144


144

Figure 5.8 – Relationship between average pull-out shear stress and pull-out displacement for tests under submerged condition (Sr≈98%)

145

Figure 5.9 – Relationship between peak pull-out shear resistance and applied overburden pressure for tests in soil at degree of saturation of 38%

145


146


146

Figure 5.12 – Relationship between peak pull-out shear resistance and applied overburden pressure for submerged tests

147

Figure 5.13 – Relationship between the peak apparent coefficient of friction and applied overburden pressure for tests in soil atdegree of saturation of 38%

147

Figure 5.14 – Peak apparent coefficient of friction vs. applied overburden pressure for tests in soil at degrees of saturation of (a) 50%, (b) 75% and (c) 98%

148

Figure 5.15 – Relationship between measured strain and pull-out displacement during pull-out for test with overburdenpressure of 40 kPa and degree of saturation of 75%

149

Figure 5.16 – Strain distribution along the nail at peak pull-out stress for test with overburden pressure of 40 kPa and degree of saturation of 75%

149

Figure 5.17 – Relationship between pull-out force and strain measured by Strain gauge 1 for test with overburden pressure of 40 kPa and degree of saturation of 75%

150

Figure 5.18 – Illustration of pull-out procedure 150

Figure 5.19 – Calculated theoretical axial tensile force vs. pull-outdisplacement

151

Figure 6.1 – (a) Effective earth pressure and (b) pore water pressure vs. time during saturating the soil – for a submerged test with overburden pressure of 200kPa

162

Figure 6.2 – Relationship between the decrease in average earth pressure immediately after grouting and applied overburden pressurefor tests at different degrees of saturation

163

Figure 6.3 – Changes of average earth pressure at different stages of testing for tests with soil at (a) 38% and (b) 50% degrees ofsaturation

164

Figure 6.4 – Changes of average earth pressure at different stages of testing for tests with soil at (a) 75% and (b) 98% degrees ofsaturation

165

Figure 6.5 – Surfaces of the drilllhole before and after pull-out in a test with soil at 38% degree of saturation

166

Figure 6.6 – Surfaces of the drillhole before and after pull-out in a test with soil at 50% degree of saturation

167


168


169

Figure 6.9 – Nail surfaces after pull-out in tests with soil at degrees of saturation of (a) 38% (b) 50% (c) 75% and (d) 98%

170

Figure 6.10 – Relationship between average pull-out shear stress andpull-out displacement for tests at different degrees of saturation under overburden pressure of 40kPa

171


171

Figure 6.12 – Relationship between average pull-out shear stress andpull-out displacement for tests at different degrees of saturation and overburden pressure of 120kPa

172

Figure 6.13 – Relationship between average pull-out shear stress andpull-out displacement for tests at different degrees of saturation and overburden pressure of 200kPa

172


173

Figure 6.15 – Relationship between peak pull-out shear resistance anddegree of saturation with overburden pressure of 40kPa

173

Figure 6.16 – Relationship between peak pull-out shear resistance anddegree of saturation with overburden pressures of (a) 80kPa and (b) 120kPa

174

Figure 6.17 – Relationship between peak pull-out shear resistance anddegree of saturation with overburden pressure of (a) 200kPa and (b) 300kPa

175

Figure 7.1 – Earth pressure vs. time during (a) drilling and grouting and (b) pressure grouting only – for a test in soil at 50% degree of saturation under overburden pressure of 200kPa and withgrouting pressure of 130kPa

185

Figure 7.2 – Variation of average total earth pressure for tests in soil at50% degree of saturation with grouting pressure of (a) 80kPaand (b) 130 kPa

186

Figure 7.3 – Failure surfaces of the soil nails for tests in soil at 50%degree of saturation with grouting pressure of (a) 0kPa (gravity head only) (b) 80kPa and (c) 130kPa

187

Figure 7.4 – Average pull-out shear stress vs. pull-out displacement for tests in soil at 50% degree of saturation with groutingpressures (GP) of 0, 80, 130kPa and under overburden pressures of (a) 80kPa and (b) 200kPa

188

Figure 7.5 – Average pull-out shear stress vs. pull-out displacement for tests in soil at 50% degree of saturation under overburdenpressures of 80kPa and 200kPa with grouting pressures of (a) 80kPa and (b) 130kPa

189

Figure 7.6 – (a) Peak shear resistance, (b) shear stress at displacement of 100mm and (c) shear stress at displacement of 200mm versus grouting pressure for tests in soil at 50% degree of saturation under overburden pressure OP=80kPa and 200kPa

190

Figure 8.1 – 3-D mesh of the finite element model 205

Figure 8.2 – Load and boundary conditions of the FE model beforepull-out – (a) front view (cross-section) and (b) side view (longitudinal section)

206

Figure 8.3 – Load and boundary conditions of the model during pull-out (side view)

207

Figure 8.4 – Vertical stress contour after drilling for a simulation with anapplied overburden pressure (OP) of 120kPa (compressivestress is negative)

207

Figure 8.5 – Vertical stress contour after pull-out for a simulation with OP=120kPa (compressive stress is negative)

208

Figure 8.6 – Maximum strain rate contour after pull-out for a simulation with OP=120kPa

208

Figure 8.7 – Vertical stress distribution before and after drilling along thepath shown in Figure 8.4

209

Figure 8.8 – Distribution of the vertical stress and increased verticalstress at peak pull-out resistance along the path shown in Figure 8.4

209

Figure 8.9 – Changes of average vertical stress at the locations of P-cells 1 to 4 at different stages of modeling under different applied overburden pressures

210

Figure 8.10 – Simulated average pull-out shear stress vs. pull-out displacement under different applied overburden pressures

210

Figure 8.11 – Comparison between the measured and simulated average pull-out shear stress- displacement curves with OP=40kPa

211

Figure 8.12 – Comparison between the measured and simulated averagepull-out shear stress- displacement curves with OP=80kPa

211


212

Figure 8.14 – Comparison between the measured and simulated averagepull-out shear stress- displacement curves with OP=200kPa

212

Figure 8.15 – Comparison the between measured and simulated average pull-out shear stress- displacement curves with OP=300kPa

213

Figure 8.16 – Simulated average pull-out shear stress vs. pull-out displacement with different dilation angles underOP=120kPa

213

Figure 8.17 – Simulated peak pull-out shear stress vs. dilation angle with OP=120kPa

214

Figure 8.18 – Simulated vertical displacement on the top surface of themodel vs. pull-out displacement with different dilation angles under OP=120kPa

214

Figure 8.19 – Procedure for simulating the pressure grouting 215

Figure 8.20 – Minimum principal stress contour after application ofgrouting pressure for a simulation with OP=80kPa andgrouting pressure (GP) of 200kPa (compressive stress isnegative)

215

Figure 8.21 – Variation of the minimum principal stress in the elements next to the hole surface throughout the simulation withOP=80kPa and GP=200kPa (compressive stress is negative)

216

Figure 8.22 – Simulated average pull-out shear stress vs. pull-out displacement under different grouting pressures with OP=80kPa

216

Figure 8.23 – Simulated average pull-out shear stress vs. pull-out displacement under different grouting pressures withOP=120kPa

217

Figure 8.24 – Simulated average pull-out shear stress vs. pull-out displacement under different grouting pressures withOP=200kPa

217

Figure 8.25 – Simulated peak pull-out shear stress vs. grouting pressure under different applied overburden pressures

218

Laboratory Pull-out Testing Study on Soil Nails in Compacted Completely Decomposed Granite Fill Su Lijun

CHAPTER 1: INTRODUCTION

1.1 BACKGROUND

Soil nailing is a technique for stabilizing soil slopes and excavations by installing a

large number of closely spaced passive inclusions (steel bar, fibre glass bar/pipe, or

other slender structures with high-tensile strength) into the in-situ soil mass. Soil nails

can be divided into several types according to the installation method, such as driven

soil nails, jacked soil nails, fired soil nails and cement (or concrete) grouted soil nails.

Cement (or concrete) grouted soil nails are the most widely used. The soil nailing

technique has been increasingly used in many countries since it originated in the early

1970’s because of its technical and economic advantages. In Hong Kong, soil nailing

has been commonly used to stabilize newly cut and sub-standard existing slopes for

decades. Every year, tens of thousands of soil nails are installed into these slopes to

increase stability.

The Hong Kong territory is mostly hilly, for 70% of the land area. Flat land is very

limited. During the rapid economic growth of the 1950’s and 1960’s, large amounts of

flat lands were required for residential, commercial, industrial and infrastructural

developments. Prior to 1977, more than 6000 cut and fill slopes were constructed at

angles of 60º or steeper, with slope heights ranging from 15 to 30m (Brand 1985). Most

of the cut slopes formed before 1977 were designed empirically to an angle of 10

vertical to 6 horizontal without much consideration of the geological or hydrological

characteristics of the slope. Slopes in Hong Kong generally consist of completely

decomposed granite (CDG) or volcanic (CDV) soil with CDG being more common.

- 1 -

Chapter 1: Introduction

Both of these two types of soils have a high strength when unsaturated during the dry

season but a low strength when fully saturated (Lumb 1975). According to the current

standard, many of those pre-1977 slopes in Hong Kong are considered to be substandard

and liable to failure in periods of intense tropical rainfall. Some publications have

reported the relationship between rainfall and landslides in Hong Kong (Brand et al.

1984, Kay and Chen 1995). A number of disasters occurred in the early and mid 1970’s

(Lumb 1975 and 1979). The two major disastrous landslides in Hong Kong’s history

occurred on 18 June 1972 at Po Shan Road and Sau Mau Ping. At the Sau Mau Ping

housing estate, a 35m high fill slope liquefied and a single-story building at the foot of

the slope was embedded (Figure 1.1), 6000 m3 of soil cascaded down and 71 were killed

and 60 injured. On the steep hillside above Po Shan Road, a slope failure with an area of

120m long and 67m wide occurred with a failure depth of 10m (about 20000 m3). This

slope failure demolished a 4-story building and a 13-story apartment block (Figure 1.2),

67 were killed and 20 injured (Government of Hong Kong 1972). There have also been

some other landslides which have caused more than 30 fatalities and more than 20

injuries. Horelli (2005) studied landslides in Hong Kong, including the major causes of

landslides, and a case study of a land slide etc. Sun (1998) carried out a review of fill

slope failures in Hong Kong.

After the disastrous landslides occurred in the early and mid 1970’s, The Hong Kong

Government started the Landslip Preventative Measures (LPM) programme. In this

programme preventative and remedial works were carried out on those prior to 1977

steeper cut and fill slopes to bring them up to current standards. The most direct method

of improving the factor of safety was to cut back the slope to a flatter angle. But in

Hong Kong’s crowded conditions, this was often not possible because of space

restrictions. In these cases, structural strengthening elements such as retaining walls,

- 2 -


cantilever caisson walls or large diameter dowels were used. From the late 1980’s, soil

nailing was found to be cheaper and more flexible in relation to the demand for space.

In some cases, this method was found to be even cheaper than the cut back method. The

use of soil nailing to stabilize existing and newly formed slopes blossomed from then on

(Powell and Watkins, 1990; Watkins and Powell 1992; Yim et al. 1988 and Forth 1997).

There are approximately 25,000 cut and fill slopes formed in hilly terrain in Hong Kong,

as identified so far by the Geotechnical Engineering Office of the Hong Kong

Government. More than HK$0.8 billion (US$103 million) is spent on slope upgrading

each year. About 80% of slopes are stabilized using soil nails. In the design of a soil

nailing system, the shear strength of the interface between a soil nail and the

surrounding soil is a key parameter for the design and stability assessment of a soil nail

stabilized slope. In current practice in Hong Kong, the interface shear strength (adhesion

ca and friction angle δ) is generally assumed to be the same as the shear strength

(cohesion c and internal friction angleφ ) of the soil itself. The interface shear strength is

then normally verified in the construction stage by field pull-out tests on soil nails

(Powell and Watkins 1990). There are considerable uncertainties about the evaluation of

interface shear strength parameter with field verification tests. For example, the shear

strength at the interface between the cement grout surface and the surrounding soil is

different from that of the soil itself. In addition, the field soil nail pull-out tests are

subjected to actual site conditions that have high local variations and the results

therefore in many cases are scattered even for two adjacent test nails within the same

site. Another problem is that the field pull-out tests are usually conducted under

conditions that are less severe than design conditions (e.g. the soil is not saturated) and

as a result the in-situ pull-out resistance may not be representative of the worst site

conditions. To enable a reliable and economic design of a soil nailing system,

- 3 -


understanding of the fundamental interaction mechanism and shear strength between a

cement grouted soil nail and the surrounding soil is of great significance.

Laboratory pull-out tests can be used to study the interaction mechanism and shear

strength between a soil nail and the surrounding soil under controllable conditions.

Many laboratory pull-out tests have been carried out by other researchers in other

countries, such as the tests conducted by Chang et al. (1996), Hong et al. (2003) and

Franzén (1998). In Hong Kong, most soil nail pull-out tests are performed on

completely decomposed granite (CDG). Laboratory pull-out tests on both steel bars and

grouted nails were carried out by Lee et al. (2001), Junaideen et al. (2004), Pradhan et

al. (2003) in the Hong Kong University using a large pull-out box in loose CDG fill.

The vertical pressure was applied by a hydraulic jack acting on a rigid plate on the top

soil surface. The stresses on the soil nail were not uniform and the water saturation

control was difficult (Pradhan et al. 2003). Chu (2003), Chu and Yin (2005, 2005a)

carried out a series of laboratory pull-out tests using a 700mm×570mm×610mm

pull-out box on grouted nails in compacted CDG fill. As a general observation from

these studies, the pull-out resistance was found to increase with the applied overburden

pressure and decrease with increasing degree of saturation of the soil. However, the

influences of the drilling process, stress release, pressure grouting etc. were not studied

at all. In addition, the box used by Chu and Yin (2005a) was a simple one without good

instrumentation. In order to overcome these disadvantages and to meet the demands for

more comprehensive studies on the soil nail pull-out resistance, an innovative pull-out

box was designed. In order to accelerate the testing programme, two boxes were

constructed and setup in the Soil Mechanics Laboratory of The Hong Kong Polytechnic

University. Each pull-out box is fully instrumented with earth pressure cells,

tensiometers, etc. inside the box. A number of pull-out tests have been carried out using

- 4 -


the two boxes to study the relationship between the pull-out resistance and certain

influencing factors, such as the overburden pressure after hole drilling and stress release,

the degree of saturation and dilatancy of the soil, and cement pressure grouting.

1.2 OBJECTIVES

The objectives of this research project are to study the influences of a few factors on the

nail-soil interface shear resistance based on laboratory pull-out tests which simulate the

procedure of soil nailing construction in the field. The following specific issues have

been studied:

(a) The pull-out resistance of soil nails in a CDG soil at the same degree of

saturation under different overburden pressures.

(b) The effect of the degree of saturation of the soil on the pull-out resistance soil

nails.

(c) The stress release effect during the drill hole procedure by monitoring the earth

pressures at locations close to the hole surface with earth pressure cells.

(d) The influence of soil dilatancy on the pull-out resistance soil nails.

(e) The influence of pressure grouting on the pull-out resistance of soil nails.

1.3 ORGANIZATION OF THE THESIS

This thesis consists of nine chapters. Chapter 1 briefly introduces the background to the

soil nailing technique and its use in Hong Kong. The objectives and methodology of this

- 5 -


research project are presented.

Chapter 2 reviews the origins and history of soil nailing. The definition of soil nailing,

the mechanism and failure modes of soil nailing systems and the fields of soil nailing

applications are presented and discussed. Different design methods are reviewed.

Current design codes and guides for soil nailing practice are examined. Research work

for developing the soil nailing technique and its analysis in the literature are reviewed

and summarized.

The design and construction of the laboratory pull-out box and relevant apparatus are

described in Chapter 3. A detailed description of the determination of the dimensions of

the box is presented. The design and details of the apparatus for saturating the soil,

drilling the hole, grouting and pull-out of the nail are described.

Chapter 4 presents the basic properties of the soil and cement grout, the calibration of

the sensors and transducers used in the pull-out tests and the test procedures. The results

of the tests for determining the properties of the soil and the cement grout are presented

and discussed. Methods and procedures for calibrating each type of sensor and

transducer are described.

Chapter 5 presents the results of tests on soil nails in a compacted CDG fill at degrees of

saturation of 38%, 75% and 98% under applied overburden pressures of 40kPa, 80kPa,

120kPa, 200kPa and 300kPa respectively. The stress release effects, the distribution of

shear stress on the nail-soil interface along the length of a nail and the influence of

overburden pressure on the soil nail pull-out behaviour and resistance are discussed.

- 6 -


Chapter 6 discusses the influence of the degree of saturation of the soil on the soil nail

pull-out behaviour and resistance. Variations of earth pressure, failure patterns and

pull-out resistance of the soil nail in tests with CDG soil at different degrees of

saturation are compared and discussed.

In Chapter 7, the results of the tests with nails grouted under grouting pressures of

80kPa and 130kPa under overburden pressures of 80kPa and 200kPa respectively in the

soil at the same degree of saturation of 50% are presented. The variations of earth

pressure and failure pattern of the soil nail in these tests are discussed. Investigation into

the influence of grouting pressure on the soil nail pull-out behaviour and resistance is

also discussed.

In Chapter 8, the establishment of a three dimensional finite element model is described.

Simulation results are presented and discussed.

Chapter 9 presents a summary of major conclusions drawn from this research project

and suggests recommendations for further research in this subject.

- 7 -


Figure 1.1 – 1972 Sau Mau Ping Landslide

- 8 -


Figure 2.2 – 1972 Po Shan Road Landslide

- 9 -

Chapter 2: Literature review

CHAPTER 2: LITERATURE REVIEW

2.1 THE SOIL NAILING TECHNIQUE

Soil nailing is a technique for stabilizing in-situ soil mass by installing a large number

of closely spaced passive slender inclusions into the soil mass. The inclusions are called

soil nails. Soil nails can be steel bars, steel strips and steel bars surrounded by cement

grout or concrete. In recent years, Fiber glass Reinforced Plastic (FRP), a material with

high tensile strength, is increasingly being used to replace steel bars in cement grouted

soil nails because of its high corrosion resistance and advantages in environmental

protection (Ortigao and Palmeira 1997). Soil nails can resist tensile stress, shear stress

and bending moments. The soil nailing technique originated as the New Austrian

Tunneling method evolved in the early 1960’s by combining shotcrete and fully bonded

steel inclusions to support a rock tunneling system. It became popular in Europe and

North America in the 1970’s. In Hong Kong, soil nailing has been widely used for

stabilizing slopes since the late 1980’s.

Soil nails are divided according to installation methods, into driven, fired (or launched),

jet-grouted, corrosion-protected and cement (or concrete) grouted soil nails. The most

widely used nail is the cement (or concrete) grouted type.

Driven soil nails, commonly used in France and Germany, are small-diameter rods or

bars, or metallic strips, made of mild steel with yield strength of 350MPa. They are

closely spaced (2 to 4 bars per square meter) and create a rather homogeneous

composite reinforced soil mass. The nails are driven into the ground at the designed

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inclination using a vibropercussion pneumatic or hydraulic hammer with no preliminary

drilling. Special nails with an axial channel can be used to allow for grout sealing of the

nail to the surrounding soil after its complete penetration. This installation technique is

rapid and economical (4 to 6 nails per hour). However, it is limited by the length of the

bars and by the inhomogeneous of the soil (e.g., the presence of boulders).

Fired (launched) soil nails are directly fired into the soil mass using a compressed air

launcher (Bridle and Myles 1991; Myles and Bridle 1992) as shown in Figure 2.1. The

nails are installed at speeds of 200mph (89.39m/s) with an energy transfer of up to

100kJ. This installation technique enables optimization of nail installation with a

minimum of site disruption. During penetration the ground around the nail is displaced

and compressed. The technology is currently used primarily for slope stabilization.

Jet-grouted nails are composite inclusions made of a grouted soil with a central steel rod.

This technique combines vibropercussion driving and high-pressure jet grouting. The

nails are installed using a high frequency vibropercussion hammer, and cement grouting

is performed during installation. The grout is injected through a small-diameter channel

in the reinforcing rod under a pressure that is sufficiently high to cause hydraulic

fracturing of the surrounding soil. However, in granular soils, nailing with a lower

grouting pressure has also succeeded. The jet-grouting installation technique provides

hydraulic fracturing and re-compaction of the surrounding soil and significantly

increases the pull-out resistance of the soil nail.

Corrosion-protected nails generally use double protection schemes similar to those

commonly used in ground anchor practice. They are usually used in permanent

structures. For permanent applications of soil nailing, based on current experience, it is

- 11 -


recommended (Elias and Juran, 1991) that a minimum grout cover of 1.5 inches

(38.1mm) is achieved along the total length of the nail for corrosion protection purpose.

Secondary protection should be provided by electro statically applied resin-bonded

epoxy on the bars with a minimum thickness of 0.35mm.

Cement grouted nails can be used in both temporary and permanent construction. They

are generally steel bars (15 to 46mm in diameter) which are placed into boreholes (100

to 150mm in diameter) and grouted with cement (or concrete) grout. The vertical and

horizontal spaces between nails vary typically from 1 to 3m depending on the type of

the in-situ soil. The nails are grouted by gravity or under low pressure. Ribbed bars can

be used to improve the nail-grout adherence. This method can provide relatively high

pull-out resistance and is widely used in Hong Kong.

2.1.1 Characteristics of soil nailing

Compared with some traditional soil reinforcement methods, soil nailing has the

following characteristic features (Bruce and Jewell 1986 and 1987a):

a) Compared with the reticulated micro piles

The reticulated micro piles are steeply inclined in the soil at various angles both

parallel and perpendicular to the wall face. The overall aim of this method, similar

to soil nailing, is to provide a stable block to support the soil mass behind. But soil

nails are generally installed into the soil mass horizontally or at small and usually

similar angles (Figure 2.2).

b) Compared with soil dowelling

Soil dowelling is applied to reduce downslope movements for relatively flat slopes

- 12 -


with a small number of large diameter dowels. Soil nailing, by contrast, uses a large

number of small diameter inclusions (Figure 2.2).

c) Compared with prestressed ground anchorages

Ground anchorages are stressed after installation and their effectiveness does not

depend on ground deformation in service. But soil nails are not prestressed and

require a finite soil deformation to mobilize their strengthening forces.

Soil nails are in contact with the soil along their whole length but ground anchorages

transfer loads only along the fixed anchorage length.

Soil nails are installed at a much higher density than ground anchorages. Thus the

failure of a single nail is not very important for the whole system.

Strong bearing facilities must be provided at the head of a ground anchorage to

resist the high prestress. But for soil nails, small steel bearing plates or simple

welding of the nail head to the rebar mesh are enough.

d) Compared with reinforced earth walls

The reinforced earth method has many similar features to the soil nailing method.

They are both horizontally installed (or nearly horizontally) into the soil mass with

high density. The reinforcement is placed in the soil unstressed and the

reinforcement forces are mobilized by subsequent deformation of the soil. The

reinforcement forces are resisted by frictional bond between the soil and the

reinforcement. But the construction sequences of these two methods are totally

different. For strengthening excavations, the soil nailing is constructed by staged

- 13 -


excavations and installation of nails from “top-down”. For reinforcing existing

slopes, the construction is completed by direct drilling and installation of nails. But

reinforced earth structures are constructed “bottom-up” (Figure 2.3).

2.1.2 Advantages and limitations of soil nailing

The most important advantage which has contributed to the growing popularity of soil

nailing is its economic advantage. Soil nailing can save 10% to 30% of the cost

comparing with the ground anchorage method (Bruce and Jewell 1986). In Hong Kong,

the direct cost of soil nailing is generally similar to that for cutting back solutions, but

since no additional land is required the overall cost is significantly lower (Powell and

Watkins 1990).

The construction equipment of soil nailing is relatively small scale, mobile and quiet.

This is an important advantage in urban environments where the construction space is

congested and noise and vibration may cause problems. The low vibration will cause

fewer disturbances to the surrounding existing buildings.

The construction of soil nailing system is flexible. For drilling, coring, rotary drilling

and even hand drilling can be used according to the site conditions. The facings of the

soil nailing system are also diversified, such as shotcrete facing, precast concrete facing

etc. For slope stabilization sometimes there is even no facing and only steel plates are

used to fix the nail heads.

Even though soil nailing has the above advantages, it still has some limitations. For

stabilizing excavations, soil nailing construction requires the formation of cuts generally

1-2m high in the soil. These cuts must stand up unsupported for at least a few hours

- 14 -


prior to shotcreting and nailing. If the soil is not strong enough, a pretreatment may be

necessary to stabilize the face.

Extra care should be taken when using soil nailing in soft clay. The low frictional

resistance of soft clay would require a very high density of reinforcements of

considerable length to ensure adequate levels of stability (Bruce and Jewell 1986).

In urban areas, a closely spaced array of reinforcements may interfere with nearby

utilities. Soils close to utility trenches may contain poorly compacted or unsuitable fill

for soil nailing.

The horizontal displacements for structures supported with soil nails may be greater

than for those supported by prestressed tiebacks. This may cause distortions to

immediately adjacent structures (Byrne et al. 1998).

2.1.3 Fields of application

Soil nailing has been successfully used in both temporary and permanent constructions

for stabilizing existing natural and fill slopes or new cut slopes and excavations (Bruce

and Jewell 1986). The fields of application for soil nailing are summarized as follows:

(a) New construction

1) Support of deep excavations for high rise buildings or underground constructions

for car parks and so on (Shen et al. 1981 and 1981a; Cheang et al. 1999).

2) Retaining walls for protecting highways or railways from being affected by

landslips (Bruce and Jewell 1987; Wehr 2003).

- 15 -


3) Stabilizing tunnel portals and tunnel facings (Ortigao et al. 1995; Barley and

Graham 1997; Ng and Lee 2002).

4) Reinforcement of bridge abutments (Hanna et al. 1998; Briaud and Lim 1997).

5) Stabilizing cut slopes (Pedley and Pugh 1995).

(b) Remedial works

Remedial works include the strengthening of existing marginally stable natural or fill

slopes and existing old retaining structures (Bruce and Jewell 1986; Schwing and

Gudehus 1988). The types of projects in this category include:

1) Repair of masonry gravity retaining walls after or before failure to prevent

possible movements behind the wall.

2) Repair of reinforced earth walls to replace the effect of the reinforcing strips or

fasteners damaged by overloading or corrosion.

3) Reinforcement of failed slopes due to failure or inadequate strength of preexisting

supports. Stabilization of marginally stable slopes to eliminate any threat to other

nearby constructions (Guilloux and Schlosser 1982; Cali 1996).

4) Stabilization of anchored walls after failure of the prestressed rock anchorages

caused by structure overloading or by corrosion of tendons.

- 16 -


2.1.4 Soils suitable for soil nailing

For stabilizing excavations, the ideal soil for soil nailing should be able to stand vertical

for a height of 1-2 meters for 1-2 days. In addition, the soil nails should be able to be

easily installed, that is, there should not be large amount of cobbles and boulders in the

soil. At the same time, the soil must provide enough bond or frictional resistance on the

interface between soil nails and the surrounding soil. Soil types that are ideal for soil

nailing include glacial till with a limited number of boulders and cobbles, weathered

rock (such as CDG), cemented sands, stiff silts and dense sands and silts. Applications

of soil nailing in residual soil (weathered rock) have been reported by Khalil et al. (1998)

and Sigourney (1996). Successful applications of soil nailing in loess and moraine soils

have been presented by Guilloux et al. (1983) and Ho et al. (1989) respectively.

Soil nailing is normally not recommended for use in plastic clays, clean granular soils

and varved and organic silts, especially for permanent constructions. For these types of

soils, the nail-soil interface shear resistances are very small and the nail needs to be

excessively long to reduce the deformation of the structure. In addition, clean granular

soils, varved and organic silts generally can not remain vertical for the time required to

install the nails and shotcrete. In order to provide more data on soil nailing in these

types of soils, Oral and Sheahan (1998) and Sheahan (2000) studied an experimental

soil nailing wall in soft clay. Based on the observation of the deformation and failure

mode of the experimental wall, they concluded that soil nailing can be effectively used

for temporary excavation support in clayey soils.

2.2 BEHAVIOUR OF SOIL NAILING

2.2.1 Soil nailing mechanism

The fundamental mechanism of soil nailing is the development of tensile forces in the

- 17 -


passive reinforcements as they react against the lateral deformations of the structure. In

the case of a top-down constructed soil nail wall, the lateral expansion of the reinforced

zone is associated with further excavations. In the case of repair of existing retaining

structures or the stabilization of marginally stable slopes, the lateral deformations are

due to the movements of the wall or slope as a result of inadequate support. In both

cases, the reinforcements interact with the ground to support the stresses and strains that

would otherwise cause the unreinforced ground to fail. These reinforcements are

oriented to correspond in general with the direction of maximum tensile straining within

the soil so that the generation of tensile forces is dominant (Byrne et al. 1998).

The tensile forces are developed within the soil nails primarily as a result of the friction

interaction between the nail and the surrounding soils, and secondarily by the

soil-structure interaction between the nail head and the soils. The tensile force at the nail

head is much smaller than the maximum tensile force in the nail. The maximum tensile

forces are located at the intersections of the nails and the potential failure surface. This

surface divides the reinforced soil mass into two zones as shown in Figure 2.4:

An “active zone” located immediately behind the facing, where the frictional shear

stresses exerted on the surface of the nail are oriented towards the facing and tends to

pull out the nails.

A “passive zone” located behind the potential failure surface, where the frictional shear

stresses are directed towards the inside of the soil nail structure and tends to restrain the

reinforcements from pull-out (Schlosser 1982; Guilloux and Schlosser 1982).

- 18 -


2.2.2 Nail-soil interface shear resistance

As discussed above, the development of shear stresses on the interface between soil

nails and the surrounding soil is the fundamental interaction mechanism between the

nail and the ground. Thus the nail-soil interface resistance becomes a key parameter

which controls the design, deformation and stability assessment of soil nailed structures.

In order to fully utilize the tensile strength of the steel bar, the pull-out failure should be

avoided in soil nailing design and therefore the pull-out resistance provided by that part

of the nail inside the passive zone should be sufficiently large.

The nail-soil interface shear resistance is related to many factors, such as the normal

stress exerted on the nail surface, the shear strength of the soil, the roughness of the nail

surface, the nail perimeter, soil dilatancy and etc. Since the 1980’s, many authors have

studied the soil nail pull-out resistance by analytical and empirical methods including

Schlosser and Guilloux (1981), Cartier and Gigan (1983), Jewell (1990), HA 68/94

(1994) and Luo (2002). In these studies the parameters involved were normal stress

acting on the nail surface, shear strength of the soil, nail perimeter and soil dilatancy.

These methods are classified into two categories according to whether the soil nail

pull-out resistance is dependent on the depth or not.

Schlosser and Guilloux (1981) developed an equation for calculation of the ultimate

pull-out resistance:

*2 μσ vef DcPT ′+′= (2-1)

where Tf is the pull-out force per linear meter; P is the perimeter of the nail; c′ is the

effective cohesion of the soil; De is the width of an equivalent flat reinforcement strip;

vσ ′ is the theoretical vertical stress calculated at the mid-depth of the reinforcement;

- 19 -


*μ is the coefficient of apparent friction of the soil.

The apparent friction coefficient is defined by dividing the maximum shear stress *μ

maxτ by the theoretical effective vertical stress (Schlosser 1982):

vστ

μ′

= max* (2-2)

Schlosser and Guilloux (1981) observed that the apparent friction coefficient

decreased with depth and the reason might be the decrease of dilatancy. The

decreased with depth and became equal to

*μ

*μ

φ′tan (φ′ is the effective internal friction

angle of the soil) below a certain depth. The pull-out resistance was considered to be

independent of the depth because the decrease of the apparent friction coefficient

was compensated by the increase of the theoretical effective vertical stress. This was

confirmed by Cartier and Gigan (1983) by pull-out tests on driven metal profiles used as

reinforcements in a nailed soil retaining wall built in silty fine sand (Figure 2.5).

*μ

This equation was adopted by many other authors directly (such as Juran 1985) or after

some modification (such as Milligan and Tei (1998) who replaced vσ ′ by mσ which is

the normal stress acting on the nail surface by considering the K0).

Jewell (1990) proposed an equation in which the pull-out capacity was related to the

average normal effective stress exerted on the soil nail surface:

φσπ tanbrapull fDLP ′= (2-3)

where Ppull is the pull-out force; La is the anchorage length of the nail; fb is the bond

coefficient; φ is the internal friction angle of the soil; rσ ′ is the average normal

effective stress acting on the circumference of the reinforcement.

- 20 -


Considering the at rest lateral earth pressure coefficient K0, the average normal effective

stress acting on the soil nail surface can be derived as:

vrK

σσ ′+=′

2)1( 0 (2-4)

Jewell (1990) pointed out that the average normal effective stress is often in the range of

7.0/1 ≥′′≥ vr σσ in steep slopes with lightly over consolidated soils. The range of fb is

from fb=1 for a fully rough surface which might be achieved at a grout to soil surface to

fb=0.2 to 0.4 for a smooth surface as might be apply between soil and a smooth metal

surface.

In HA 68/94 (1994) a similar equation which considered the cohesion of the soil was

suggested:

)tan( φσλ ′′+′= neppull cLP

(2-5)

where Le is the anchorage length of the nail; c′ and φ′ are the effective cohesion and

internal friction angle of the soil; nσ ′ is the average normal effective stress acting on

the nail surface and pλ is the pull-out factor:

hholep Sd /απλ = (2-6)

where dhole is the diameter of the borehole; Sh is the horizontal spacing of the nails and

α is the interface sliding factor which is similar to the bond coefficient fb and is

defined as:

)tan()tan( intint

desdesv

v

cc′+′′′+′′

=φσφσ

α (2-7)

where the subscripts int and des relate to interface and design values respectively.

- 21 -


Luo et al. (2000 and 2002) developed a theoretical model, in which the soil dilatancy is

considered, to calculate the pull-out resistance. In this model the apparent friction

coefficient for the peak state and critical state are given by:

)tan( maxψφσ

η +′′

′= cv

v

pp

q (2-8)

cvv

cvcv

qφ

ση ′

′′

= tan (2-9)

where and are the average normal stresses on the soil nail for the peak and

critical states respectively;

pq′ cvq′

cvφ′ and maxψ are the effective internal friction angle at the

critical state and the maximum dilation angle of the soil.

This model was verified by simulating the pull-out tests presented by Schlosser et al.

(1992) and Luo (2001) and good agreement between the test results and predicted

results was achieved.

2.2.3 Influence of bending stiffness of the nail

The soil nails may be subjected to shear forces and bending moments along the potential

failure surface in the soil nail structures (Figure 2.6). Many authors have investigated

the contribution of the bending stiffness of the nails to the stability of a soil nail

structure, such as Schlosser (1982), Jewell (1990), Jewell and Pedley (1990 and 1992)

and Juran et al. (1990). But different authors have different points of view on the

influence of the bending stiffness of the nail. Schlosser (1991) claimed the effect of the

bending stiffness can be either beneficial or not beneficial depending on the behaviour

of the soil nail system. Juran et al. (1991) concluded that increasing the nail bending

stiffness could result in a decrease of structure stability. But Jewell and Pedley (1990

and 1992) thought otherwise and considered that the bending stiffness of the nail is only

- 22 -


of secondary importance compared with the tensile resistance and can be ignored in soil

nailing design.

2.2.4 Failure modes of soil nailed structures

Soil nailed structures include the soil nails, the facing and soils between them and they

work together to support the soil mass behind them. Soil nailed structures may fail due

to internal or external failure.

Internal failure is failure which occurs inside the reinforced zone and can be considered

as one or a combination of the following:

(a) Tension failure of the nail: The rupture of the steel bar may occur if the steel bar is

not strong enough but the bond strength between the nails and the surrounding soils

in the passive zone is large enough to ensure the nails can not be pulled out.

(b) Pull-out failure: In contrast, when the steel bar is strong enough but the bond

strength between the nails and the surrounding soils in the passive zone is not large

enough, pull-out failure may occur.

(c) Facing failure: When the bond strength in the passive zone and the tensile strength

of the steel bar are both large enough but the bond strength in the active zone and

strength of the facing are not large enough, head or facing failure may occur.

External failure is defined as failure which occurs behind the reinforced zone. In this

type of failure, the soil nailed structure may fail like any retaining structure would, such

as by overturning, sliding, insufficient bearing capacity and overall slope failure. The

main reason for this type of failure is insufficient length of the nails together with poor

- 23 -


quality foundation etc. In this type of failure, the soil nailed structure fails as a

monolithic block and the reinforcements have not been brought into play. Thus in the

design of soil nailing, external failure should be avoided. Combination of soil nailing

and other stabilizing methods such as prestressed tiebacks is a good way of avoiding

this problem. Kim et al. (1995) studied the failure mechanism of soil nailed structures

under surcharge loading using reduced model tests. Cardoso and Fernandes (1991)

investigated the failure mechanism of soil nailed walls using the Finite Element

Method.

2.3 DESIGN METHODS FOR SOIL NAILING STRUCTURES

There are several different methods currently used for the design of soil nailed

structures. All these methods are based on limit equilibrium analysis in which the

potential failure surfaces throughout the soil mass are examined. Both internal and

external failures are examined in these methods. Most of these methods are based on the

slope stability concept and are the same as the limit equilibrium analysis methods for

slope stability except that the resistance of the soil nails at the failure surface are also

included. Famous methods of this type are the Schlosser method (1982) in France and

Davis (Shen et al. 1981) method in the USA. The former method considers the tensile

resistance as well as the shear resistance of the reinforcement in the analysis, while the

later method uses only the tensile resistance. There are also some methods that consider

the soil nailing support system as a composite retaining wall when conducting a stability

analysis and design, such as the Stoker and Gasler (1979) method in Germany. All the

above are overall stability analysis methods and can not assess the internal force in each

soil nail. Juran et al. (1988) developed a limit equilibrium method using logarithmic

spirals failure surface. In this method, an empirical earth pressure distribution is given

and soil nail forces can be estimated based on this distribution. These methods are

- 24 -


briefly presented in the following paragraphs.

2.3.1 The Davis method

The Davis method was developed at the University of California at Davis, by Shen et al.

(1981b). A parabolic curve passing through the toe of the wall is assumed to represent

the failure surface for an in-situ reinforced soil mass. A classical method-of-slices slope

stability analysis is used to evaluate the contribution of the nails to overall stability.

Only tensile forces are considered to be mobilized in the reinforcements. The tensile

forces are divided into components parallel and perpendicular to the failure surface. The

normal and tangential components of the tensile forces in all of the reinforcements

crossing the failure surface are added to the resisting forces mobilized in the soil when

determining the factor of safety of the entire mass. To carry out the stability analysis,

two conditions must be considered separately. One condition is that the failure surface

must be entirely within the reinforced soil mass and the other condition allows the

failure surface to extend beyond the reinforced zone (Figure 2.7). As with conventional

slope stability analysis, a factor of safety for the soil nail reinforced soil mass can be

obtained by iteration.

Limit analyses performed in accordance with the above method have been compared

with the finite element analysis results carried out by Bang (1979) on an in-situ

reinforced soil excavation. Figure 2.8 shows the predicted potential failure surfaces

using the two methods and the results are in good agreement. Bang (1992) further

introduced this method and carried out some parametric studies. Bang (1996) developed

an approximate solution to evaluate the factor of safety for soil nailed walls against deep

seated failure based on this method.

- 25 -


2.3.2 The French method

The French method developed by Schlosser (1982) considers the tensile resistance,

shearing strength and bending stiffness of the nails, which contribute to the overall

stability of the in-situ reinforced soil mass. This method is also derived from the slices

methods (Fellinius, Bishop, etc.) used in slope stability analysis. The global equilibrium

of a partly or completely reinforced soil mass between the external surface and the

potential failure surface is considered. In this method, the reinforced soil is considered

as a composite material and the failure for each inclusion and the interaction between

the soil and the inclusions are taken into account (Figure 2.9). The analysis procedure is

similar to the Davis method, however, four failure criteria are considered: shear strength

of the soil, soil-inclusion friction, soil-inclusion normal pressure and strength of the

inclusion.

The strength of the inclusions combined with the tension and shear stresses (the bending

moment is related to the shear stress) is given by means of maximum plastic work under

Tresca’s yield criterion:

122

≤⎟⎟⎠

⎞⎜⎜⎝

⎛+⎟⎟

⎠

⎞⎜⎜⎝

⎛

c

c

n

n

RT

RT (2-10)

where Tn and Tc are the tension and shear force respectively and Rn and Rc are their

maximum allowable values. The corresponding yield curve is an ellipse in the (Tc, Tn )

plane. The shear strength of the soil is determined by Mohr Coulomb’s failure criterion

and the soil-inclusion friction is governed by the Schlosser and Guilloux (1981)

criterion which is discussed above. For the soil-inclusion normal pressure, considering

the combination of failure within the soil and the inclusions, the final yielding curve is

shown in Figure 2.10.

- 26 -


A particular factor of safety is used for each failure criterion. For the strength of the

inclusion, a factor of safety of 1 is used. A factor of safety of 2 is applied for the

soil-inclusion normal pressure. For the shear strength of soil, a minimum factor of

safety of 1.5 is generally required relative to overall slope stability. For the

soil-inclusion friction, the factor of safety is taken as equal to that of the soil shear

strength and has to be equal to or greater than 1.5.

2.3.3 The German method

Stocker et al. (1979) proposed a limit equilibrium approach (German method) for

designing soil nailed structures using a bilinear failure surface and assuming that nails

can withstand only tensile force. The inclinations of the two-part wedge passing through

the toe are determined iteratively to obtain a minimum factor of safety (Figure 2.11). A

force polygon is constructed by considering the forces acting on a rigid soil wedge

limited by the potential failure surface. The soil is assumed to be homogeneous and

without water. For the design of permanent structures a minimum global factor of safety

of 1.5 is recommended, the suggested partial factor of safety for soil strength parameters

is 1.2 to 1.25 and a factor of safety for friction of the inclusions is in the range of 1.5 to

2.0.

Further investigations of this method were reported by Gassler and Gudehus (1981),

Gassler and Gudehus (1983) and Gassler (1988). These studies indicated that the

bilinear failure surface is only suitable for nearly vertical nailed walls in cohesionless

soils with nails of constant length or shorter nails in the upper rows subjected to high

surcharge loads. For less steep walls or longer nails in the upper rows the circular failure

surface is more suitable. Later observations on the behaviour of soil nailing structures

which subject to self-weight of the soils and the nails themselves also shows that the

- 27 -


bilinear failure surface is not applicable for this circumstance.

2.3.4 The Juran method

Juran et al. (1988 and 1990) proposed a limit equilibrium analysis method — the

kinematical limit analysis design approach which can estimate the maximum tension

and shear forces mobilized in each inclusion. The main design assumptions (Figure 2.12)

are:

(a) The potential failure surface is a log-spiral intersecting the bottom of the wall.

(b) At failure, the locus of maximum tension and shear forces coincides with the failure

surface.

(c) The quasi-rigid active and resistant zones are separated by a thin layer of soil at a

limit state of rigid plastic flow.

(d) The shear strength of the soil is entirely mobilized all along the failure surface.

(e) The horizontal components of the interslice forces acting on both sides of a slice

containing a nail are equal.

The reinforced soil mass is divided into slices parallel to the nails with a single nail row

in one slice. The tensile and shear forces developed in each nail row at their point of

intersection with the failure surface are determined by considering the local equilibrium

in each slice. The soil nailing structure can be designed to prevent failure by pull-out

and breakage of reinforcements. This approach can evaluate the effect of the main

- 28 -


design parameters (inclination, bending stiffness and spacing of the nails) and the

structure geometry on the stability of the structure. The method was verified by

comparing the predicted results with measured results in laboratory models and

full-scale structures and the agreement between them is fairly good.

Besides proposing the kinematical analysis design method, Juran also conducted much

other work related to soil nails. Juran et al. (1982) studied the soil-bar interaction

mechanism in soil nailing by experimental and numerical methods. Juran (1985)

compared soil nailing and reinforced earth retaining structures, summarized the main

results of both laboratory model studies and full scale experiments and illustrated the

mechanism and proposed design methods for soil nail structures. Juran and Elias (1990)

presented some field observations during and after construction on instrumented full

scale structures and compared the measured results with results predicted by Juran

method. Juran et al. (1990a) compared the foregoing design methods (Table 2.1),

recommended a design procedure and provided useful charts for preliminary design.

There remains other work done by Juran which is not practical to be listed here one by

one.

Many other design methods existed. Briddle (1989) and Briddle and Barr (1990)

developed a limit equilibrium design method using a log-spiral failure surface and

dividing the active zone into vertical slices. Sabhahit et al. (1995) proposed a design

method by calculating the total reinforcement force required to raise the factor of safety

to a desired value for an unreinforced slope using Janbu’s rigorous method. Patra and

Basudhar (1997 and 2005) improved this method by further considering the location,

length, diameter and orientations of nails and considering not only overall equilibrium

but also internal equilibrium. Yuan et al. (2003) introduced a design method based on

- 29 -


the method of slices and reliability analysis. Sheahan and Ho (2003) proposed a

simplified trial wedge method which is preferable when commercial software for

traditional methods is not available.

2.3.5 Discussion on current design guides and codes

There is no universally accepted document which provides definitive guidance on soil

nail design. But there are a number of published documents which provide direction and

advice, including the UK Highways Agency’s advice note HA 68/94, the British

Standard BS 8006: 1995 and the US Department of Transportation’s design manual

named “Manual for Design & Construction Monitoring of Soil Nail Walls”. These

documents are intended to ensure good design practices and are based on the knowledge

of experts in this field and previous research and projects undertaken.

The HA 68/94 named “Design methods for the reinforcement of highway slopes by

reinforced soil and soil nailing techniques” was originally prepared to give guidance on

the design requirements for the strengthening of highway earthworks using reinforced

soil and soil nailing techniques. A single design method based on limit equilibrium of a

two-part wedge failure mechanism is offered that applies equally to both reinforced soil

and soil nailing (Love and Milligan 1995). The constraints on the mechanism are that

the inter-wedge boundary should be vertical, and that the base of the lower wedge

should intersect the toe of the slope. The two part wedge is a slightly conservative

approach for the analysis of slopes steeper than 60°. However, for shallow slopes which

are less than 27° it can be overly conservative and a circular failure surface might be

more applicable.

For the limit equilibrium calculation, a set of driving forces are assumed to be in

- 30 -


equilibrium with a set of resisting forces. The driving forces include the self weight of

the soil plus any surcharge load and unfactored values are used. The resisting forces are

represented by the shear strength of the soil and reinforcement force for which factored

design values are used. For the shear strength of the soil, the critical state parameters

cvφ′ and obtained from large displacement shear box tests or drained triaxial tests

can be directly used for granular soils or cohesive soils with PI<25%. For these types of

soils, a factor of safety ranged between 1.3 and 1.5 is recommended to be used if the

peak strength parameters

cvc′

pkφ′ and pkc′ are used. At the same time, the design value

is assumed not to be greater than 5kN/mdesc′ 2. For cohesive soils with PI>25%, cvφ′

and or the residual strength cvc′ rφ′ are recommended to be used and the design value

would normally be zero. A series of partial factors of safety is applied to the yield

strength of a nail taking account of mechanical damage, environmental effects and other

uncertainties in material strength. The nail-soil interface shear strength recommended

by HA 68/94 is discussed in the above paragraphs. HA 68/94 does not provide guidance

on serviceability limit states (Johnson et al. 2002).

desc′

The British Standard BS 8006 “Code of practice for Strengthened/reinforced soils and

other fills”, provides detailed design methods and procedures for reinforced earth fill

structures with limited advices on the design of soil nail structures. Limit equilibrium

analysis based on the two-part wedge and log-spiral failure surfaces is recommended for

the internal stability analysis of soil nail slopes. Force equilibrium is used for the

two-part wedge failure surface and moment equilibrium is used for the log-spiral failure

surface. The two-part wedge analysis is recommended because of its relative simplicity

although it may be over conservative for shallow slopes. Axial tensile forces are

considered to be the predominant stabilizing effect of the reinforcements. External

- 31 -


stability and serviceability limits governing the internal stability of soil nailed slopes are

considered to be similar to those for reinforced fill slopes.

The American Department of Transportation’s design manual provides a guideline

procedure to ensure that agencies adopting soil nail wall design and construction follow

a safe, rational procedure from site investigation through construction. Detailed design

guidance including considerations of nail and nail head strength, corrosion protection

and drainage is provided. The recommended design method has the following

characteristics:

(a) Based on slip surface limit equilibrium concepts.

(b) Considers the strength of the nail head connection to the facing, the tensile strength

of the nail and the pull-out resistance of the nail-ground interface.

(c) Provides an approach for determining the strength of the facing and the nail-facing

connection system.

(d) Recommends design earth pressures for the facing and nail head system based on

soil-structure interaction considerations and monitoring of in-service structures

(e) Introduces both Service Load Design (SLD) and Load and Resistance Factor

Design (LRFD) approaches. The allowable loads and load factors and design

strengths for the soil and reinforcements are provided respectively for SLD and

LRFD method.

(f) Recommends procedures for ensuring a proper distribution of reinforcements

within the reinforced block to enhance stability and limit wall deformation.

(g) Checks both the ultimate limit and serviceability limit.

In Hong Kong, the design of soil nailing is based on some published papers, including

- 32 -


Schlosser and Guilloux (1981), Powell and Watkins (1990) and Jewell (1990). Janbu’s

Simplified and Rigorous method are normally used to calculate the total horizontal force

required to maintain the required factor of safety. The soil nails are then designed to

provide this force (Shiu et al. 1997). The factor of safety for global failure is

recommended to be 1.4. For checking against the pull-out failure of soil nails from the

resistant zone, a minimum factor of safety of 2.0 should be applied to the ultimate

grout-steel bond strength and soil-grout bond strength to obtain the allowable design

strength. For checking against tensile failure of reinforcement, allowable tensile stresses

in the steel bars should be restricted to 55% of the characteristic strength of the steel or

230 Mpa (use the smaller one) (GEO 2005; Burland 2002).

In Hong Kong, a commercial program "Slope/W" has been widely used to aid the

design of soil nail structures. Slope/W uses the equilibrium of forces and moments to

compute the factor of safety.

2.4 FACTORS INFLUENCING THE PULL-OUT RESISTANCE

There are a number of factors influencing the pull-out resistance of a soil nail, such as

the stress conditions, methods of installation, soil conditions and soil nail surface

condition.

2.4.1 Soil conditions

The soil conditions that will influence the pull-out resistance of a soil nail include the

strength, particle size, dilatancy and degree of saturation of the soil.

One of the most important factors that influence the pull-out resistance of a soil nail is

the type of soil surrounding the nail. For example, the same type of soil nail, installed in

- 33 -


the same way into silty clay, sand and sandy gravel may give pull-out resistance values

of about 40-80kPa, 100kPa and 200kPa respectively (Bruce and Jewell 1987). The

particle size and shape of particles which are related to the soil dilatancy will

significantly influence the pull-out behaviour and pull-out resistance of a soil nail. For

soils that are composed of larger and more uniform particles with irregular shapes, the

particles will rotate and rearrange during pull-out which will result in dilation of the soil.

If the soil dilation is constrained, an increase of normal stress will result and the pull-out

resistance will increase. Luo et al. (2000) studied the influence of soil dilatancy on the

pull-out resistance. A moderate degree of water saturation of the soil is beneficial for

soil nail pull-out resistance. Soil which is either too dry or too wet is not good. Chu and

Yin (2005a) and Pradhan (2003) studied the influence of the degree of soil water

saturation on soil nail pull-out resistance and low resistance was observed for nails in

approximately saturated soils.

2.4.2 Stress conditions

The stress conditions that will influence the pull-out resistance of a nail normally refer

to the normal stress acting on the nail surface. Some authors consider that the normal

stress is related to the soil overburden pressure corresponding to depth, such as Jewell

(1990). Some other authors considered that the normal stress is independent of soil

depth, such as Schlosser and Guilloux (1981) and Cartier and Gigan (1983).

2.4.3 Methods of installation

Nails installed by different methods will have different pull-out resistances. Normally

the pull-out resistances of cement (concrete) grouted nails and jet-grouted nails are

larger than those for driven, jacked and launched nails. For driven and jacked nails, the

pull-out resistance of the former is found to be 50% greater than the latter (Franzen

- 34 -


1998). For a grouted nail, the grouting pressure will also influence the pull-out

resistance (Milligan et al. 1997).

2.4.4 The nail surface conditions

The roughness of the nail surface is an important parameter that influences the soil nail

pull-out resistance. Tei (1993) and Milligan and Tei (1998) carried out a series of

pull-out tests on both smooth and rough nails in sands. The typical displacements

needed to mobilize the peak pull-out resistance for smooth nails were found to be half

of those for rough nails. For smooth nails the pull-out resistance was observed to be

much smaller than that of rough nails under the same test conditions. The diameter of

the nail compared with the particle size of the soil was also found to have some

influence on the pull-out resistance.

2.5 RESEARCH AND DEVELOPMENT

2.5.1 Large scale model tests and field monitoring

The most direct and reliable method of studying the mechanism and failure modes of

soil nail structures is to use full scale model tests. But these tests are very expensive,

time consuming and need much space to carry out. Thus only a few large scale model

tests have been carried out in the two large soil nail projects of “Bodenvernagelung” in

Germany (1975-1980) and “Clouterre” in France (1986-1991).

Totally seven soil nail walls were constructed from top-down with vertical or nearly

vertical facings in the “Bodenvernagelung” project. Three of the seven walls were built

in sand, three in layered sand and clay or silt clay, and one in a stiff heavily over

consolidated clay as shown in Figure 2.13 (Gassler 1992). The German design method

of soil nail structures is based on these tests and the later reduced scale tests in the

- 35 -


University of Karlsruhe.

Gassler (1992) presented the results of test G in Figure 2.13. The surcharge of this test

was applied by dead loads of six reinforced concrete plates which were placed on the

top surface of the nailed wall. The measured horizontal displacements were found to be

larger than the vertical displacements and the ratio between horizontal and vertical

displacements was about 5:3 at the top of the wall. Creep displacements and change of

force distribution in the nails caused by creep effects were observed. A conclusion was

drawn that the creep property of cohesive soils can play a major role in the behaviour of

a soil nail structure.

In the French national project “Clouterre” three full scale experimental soil nail walls

were constructed to study three types of failures: breakage and pull-out of the nail bar,

and failure by excessive excavation. The first nailed wall was 7m high and built of

Fontainebleau sand and loaded from the top until failure by saturation of the soil. This

wall failed by breakage of the inclusions. The second nailed wall was the same as the

first but failed by decreasing the adherence length of the inclusions. The third wall was

failed by progressively increasing the height of the unsupported excavation (Plumelle

and Schlosser 1990; Plumelle et al. 1990; Schlosser et al. 1992; Clouterre 1993).

After failure of the first test, the nails were found to have suffered large bending

deformations and some of the top nails were broken. The tensile force was found to be

the major resisting force and bending was mobilized only when large deformation had

already occurred. At failure, a crack was observed at the top of the wall and 2.5m

behind the wall facing and this was approximately equal to the theoretical value of 0.3H

used in reinforced earth structures. A failure zone was formed at failure and the

- 36 -


maximum tensile force line and bending line were found to lie within the failure zone.

The location of the observed crack was found to coincide with the point where the

maximum tensile force line intersected the top surface of the wall (Figure 2.14).

The second experimental wall was brought to failure by gradually reducing the length of

the telescopic nails (comprising nails slid into tubes). After reducing the length of the

nails to a minimum, the whole of the soil nailed mass sank 0.27m and slid along a

well-defined failure surface, which was demarcated by the nails (figure 2.15).

The third experimental wall was a 6m high wall with the top 3m nailed and the facing of

the wall supported by one 3m and three 1m high metallic panels. The wall was brought

to failure by removing the panels one by one. Firstly the 3m high panel was removed

and the nailed wall was exposed unsupported. When the first one meter panel was

removed, a one meter excessive excavation was formed and one crack was observed but

the wall remained intact. Then the second one meter panel was removed and some sand

fell down which resulted in the formation of a soil arch. This arch remained stable

during the 24 hours before removal of the last panel. After a 45mm removal of the last

panel, the failure occurred and the nailed panel subsided 1.4m but remained attached to

the nails (Figure 2.16).

In Hong Kong, a large scale field test was carried out in the Kadoorie Agricultural

Research Center of the University of Hong Kong. The slope angle was 33°. The height

and width were 43.75m and 9m respectively. Two rows of grouted nails with diameter

of 100mm were installed at a grid of 1.5m×1.5m at an inclination of 20° from the

horizontal. The slope was brought to failure by a surcharge loading of 72kPa which was

generated by placing concrete blocks on the slope crest. Monitored data showed that the

- 37 -


nail forces had greatly improved the stability of the slope (Tang and Lee 2003).

Field monitoring has been carried out to study the deformation and serviceability

behaviour of soil nail structures especially for permanent structures. Stocker and

Riedinger (1990) presented the long term monitoring results for a 15m high soil nail

wall which had been carried out over a period of 10 years. The observed horizontal

displacements were well within the limits of suitability and serviceability. Barley et al.

(1998) conducted an eighteen-month monitoring of a soil nail slope by installing strain

gauged soil nails. During the monitoring period, loads of only 3-6kN were found to

have developed in the nails. No visually detectable movement of the slope occurred

which indicated that the soil nails fulfilled their intended purpose of stabilizing the slope.

Barley et al. (1997 and 1997a) discussed the field testing of soil nailing. Wong et al.

(1997) monitored the performance of a 9m deep permanent cut in residual soil and the

results showed that the soil nail wall performed well. Landau Associates Inc. (1999)

monitored the nail force mobilization and wall deflection of nailed highway walls.

2.5.2 Laboratory testing studies

Laboratory tests, including pull-out tests, direct shear tests, centrifuge tests and small

scale tests, have been carried out to study the failure mode and mechanism of soil nail

structures. The advantages of laboratory tests are the relatively low cost and small space

requirement. The test conditions can be well controlled and large amounts of data can be

obtained by instrumentation.

2.5.2.1 Laboratory pull-out tests

The pull-out tests carried out by other authors have been mentioned in Chapter 1 and the

above paragraphs of this chapter. Some of the tests are briefly introduced in the

- 38 -


following paragraphs.

Tei (1993) carried out a series of pull-out tests on soil nails in three types of sands using

a 254mm long, 153mm wide and 202mm high pull-out box (Figure 2.17). Two types of

nails, stiff and flexible, with different lengths and diameters were tested using the

pull-out box. Stiff nails were made of mild steel with diameter from 1.0mm to 5.2mm

and flexible nails were made of a rubber tube with external diameter of 3.0mm and

internal diameter of 2.0mm. For stiff rough nails, the pull-out displacement required to

mobilize the peak pull-out force was observed to be independent of diameter, length of

the nails and the applied vertical pressure. The peak pull-out force was found to increase

with the shear strength of the soil. Based on the data measured by strain gauges, the

axial stress in the stiff nail was observed to be approximately linearly distributed along

the nail (Figure 2.18). Analytical analysis also showed that the axial stress distribution

became more linear and the shear stress distribution became more uniform as the

relative stiffness between a nail and the soil increased. For stiff smooth surface nails, the

peak pull-out forces were found to be, as expected, much smaller than those for the stiff

rough surface nails. For the flexible nails, the peak pull-out forces were found to be

smaller than those for the stiff nails and seemed to be not proportional to the length of

the nails.

At the Chalmers University of Technology, Franzén (1998) created a 2m×4m×1.5m

pull-out box for four types (angle bar, ribbed bar, expansion bolt and round steel bar)

jacked or driven nails pulled out under four different stress levels of 25, 37.5, 75 and

125kPa. The jacked nails were installed into the soil by two hydraulic jacks and the

driven nails were installed using a percussion hammer. The objective of these laboratory

tests was to study the influence of overburden pressure, relative density, surface

- 39 -


roughness and method of installation on soil nail pull-out resistance. The results showed

that the pull-out capacity in a cohesionless soil mainly depended on the roughness of the

nail surface, relative density of the soil, nail surface area and normal pressure acting on

the nail surface. The peak pull-out resistance for driven nails was found to be 50%

higher than that for jacked nails and showed more of a strain-softening behaviour. The

residual pull-out resistance, however, seemed to be independent of the method of

installation.

In Hong Kong, the first laboratory pull-out tests were performed in completely

decomposed granite (CDG) by Lee et al. (2001) in The University of Hong Kong. Since

then, the Government, researchers and contractors showed their interest in pull-out tests

of this type of soil and more tests have been conducted. Some representative examples

are the tests conducted by Pradhan (2003) in Loosely Compacted Silty and Gravelly

Sand Fills and interface strength study tests in CDG soil by Chu (2003). The box used

by Pradhan (2003) was the 2m×1.6m×1.4m box introduced by Lee et al. (2001) and

Junaideen et al. (2004) (Figure 2.19). The interface friction angle was found to be very

close to the soil friction angle. Pull-out tests in both natural moisture content and nearly

saturated soil were carried out and it was found that the interface friction angle obtained

was similar but the cohesion reduced with a high degree of saturation. Chu (2003)

carried out a series of pull-out tests using a 0.6m×0.6m×0.7m steel box in CDG soil

which was compacted to 95% of the maximum dry density (Figure 2.20). The drilling of

the hole and installation of the nail occurred just after compaction and the vertical

pressure was applied after the cement grout had hardened. Pull-out tests in soil with

different applied vertical pressures and degrees of saturation were carried out. The

results showed that the pull-out resistance increases with the applied vertical pressure

and decreases with the degree of saturation from the natural wet to nearly saturated

- 40 -


conditions.

Smith (1992) discussed the effects of variable geology on soil nail pull-out resistance in

Hong Kong.

2.5.2.2 Direct shear and interface shear testing studies

Direct shear tests have been carried out to study the influence of the inclination, shear

and bending resistance of the inclusions on the overall strength of the reinforced soil.

Many authors have carried out this type of tests, such as Jewell and Wroth (1987),

Marchal (1990), Pedley (1990), Barr et al. (1990), Davies and Masurier (1997) and

Briddle and Davies (1997).

Jewell and Wroth (1987) carried out some direct shear tests on unreinforced soils and

soils with extensible and inextensible reinforcements. It was found that the

reinforcement has no influence on shearing resistance until the mobilized shearing

resistance in the sand exceeds the critical state value for the first time. The radiographic

observations show that the reinforcement caused more sand to deform and helped to

resist localized shear deformation. The effects of the reinforcement increase with the

stiffness of the inclusions. The maximum improvement in shear strength was achieved

when the inclination angle was about 30° from vertical (Figure 2.21). Similar tests were

carried out by Pedley (1990) using a large shear box with the internal dimension of

1m×1m×1m. It was found that surface roughness and reinforcement orientation

significantly influences the axial stress on the reinforcement but has a limited effect on

the shear stress. The shear strength improvement of the soil increases with the increase

in the stiffness and strength of the reinforcement.

Interface shear tests can be used to investigate the behaviour of soil-structure interfaces,

- 41 -


including the pile-soil interface, nail-soil interface etc. Ramsey et al. (1998) performed a

review of soil-steel interface testing with the ring shear apparatus. Chu (2003) carried out a

series of direct shear tests on the interface between a CDG soil and cement grout to

investigate soil-grout interface shearing behaviour.

2.5.2.3 Centrifuge Modeling

Centrifuge modeling is a method for studying soil structures, such as foundations,

tunnels and reinforced slopes, using greatly reduced scale models under high gravity

accelerations to represent the prototype structures under normal gravity. In the Hong

Kong University of Science and Technology, McVay et al. (1998), Zhang and McVay

(1998) and Zhang et al. (1999) carried out a series of centrifuge modeling tests to study

laterally loaded piles and pile foundations. Kamata and Mashimo (2003) carried out

centrifuge model tests of tunnel face reinforcement by bolting. This technique has been

introduced to study soil nail slopes since the early 1990’s. There are normally two types

of centrifuge modeling for soil nailing, one type is to reproduce the construction

procedure and study the behaviour and the failure of a soil nail structure under

surcharge loading (Davies et al. 1997; Aminfar 1998); another type is applying a cyclic

shaking load to the centrifuge model to simulate the behaviour of a soil nail structure

during earthquakes (Choukeir 1996; Choukeir et al. 1997; Tufenkjian and Vucetic 1992

and 2000; Vucetic et al. 1993).

Davies et al. (1997) and Aminfar (1998) carried out series of centrifuge tests on soil nail

slopes. The nailed model slope was first constructed by reproducing the construction

procedure in a number of sequential centrifuge runs. Then a surcharge load was applied

to bring the model slope to failure. The measured development of lateral displacements

of the model slope with the excavation stages was similar to that measured from the

- 42 -


actual structures. Nail axial forces and bending moments were also measured during the

tests.

Choukeir (1996) carried out series of centrifuge model tests on soil nail walls under

cyclic shaking loads and developed a seismic design method for soil nail structures on

the basis of the test results. The proposed design method was a quasi-static method

which could evaluate the seismic loading effects on the magnitude and location of the

maximum force in each inclusion. Tufenkjian and Vucetic (1992 and 2000) conducted a

number of dynamic centrifuge tests on soil nail excavation models to investigate the

stability and failure mechanisms under earthquake loading. A nearly bilinear failure

surface was observed (Figure 2.22). The nail length was observed to have a strong

influence on the stability but do not affect the failure mechanism. All these tests showed

that soil nail structures provide excellent stability under seismic loading.

2.5.2.4 Small scale tests

The first small scale model tests were conducted at University of Karlsruhe, Germany as

a part of the “Bodenvernagelung” Project. The model was 1.1m in length, 0.56m in

width and 0.72m in height and the sides of the container were constructed of Perspex to

observe the failure mode of the model. A bilinear failure surface was observed in the

nailed wall by applying surcharge to the reinforced zone.

Kitamura et al. (1988) carried out a series of reduced scale model tests for both

reinforced and unreinforced slopes. The model was 0.9m in width, 0.75m in height and

2.1m in length and the load was applied by a hydraulic jack against a loading frame

(Figure 2.23). An approximately linear and shallow failure surface was observed for

unreinforced slope. The failure surface became circular and moved backwards when the

- 43 -


inclination of the reinforcement varies in the order of upward, downward and horizontal.

The settlements of the upward and downward reinforcement at yield stress were nearly

twice as large as that of the horizontal reinforcement. This meant the horizontal

reinforcement provided the best reinforcement effects. It was observed that bending and

shear resistance of reinforcements contributed little to the reinforcement effects.

A relative large model test with a 2.0m high excavation in a 2.0m wide, 3.0m high and

6.0m long model box was carried out by Kim et al. (1996) (Figure 2.24). The loads

were also applied by hydraulic jack. The excavation was constructed by top-down

sequence to simulate the actual construction procedure. The maximum strain line

showed that the failure surface of this experimental soil nailed wall could be represented

by bilinear or log-spiral lines.

Larger model tests were carried out by Raju et al. (1997) in the Nanyang Technological

University in Singapore. Six 3.0m long, 3.0m wide and 2.4m high nailed wall were

constructed in sand with different nail lengths, nail inclinations and methods of nail

installation. The nailed wall was constructed by the top-down method and the nails were

jacked into the wall by hydraulic jack at a horizontal spacing of 1.0m and vertical

spacing of 0.5m. A uniform flexible load provided by building up a sand fill on the top

surface of the wall was applied to bring the test wall to failure. The experimental nailed

wall was observed to rotate around the toe at failure. The maximum settlements were

situated immediately behind the facing and decreased with distance from the facing.

2.5.3 Numerical modeling

Numerical modeling can be used to study nail-soil interaction, stress distribution in both

soil and nails, deformation and serviceability of soil nail structures and for conducting

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parametric studies. Most numerical modeling has been carried out to study the overall

and localized behaviour of soil nail structures, such as the investigations conducted by

Benhamida et al. (1997 and 1997a), Kim et al. (1997), Smith and Su (1997) and Cheuk

et al. (2005). Some numerical modeling has been carried out to study nail-soil

interaction behaviour, such as the works performed by Ochiai et al. (1997) and Tsai et al.

(2005). Numerical modeling can provide the stress and strain distribution throughout the

whole model which is not possible with limit equilibrium methods. But when carrying

out numerical modeling of soil nailing, one must face up to the challenge of how to

simulate the nail-soil interface behaviour.

There are normally three methods to simulate nail-soil interface behaviour. The first

method is to simulate the nail-soil interface using zero thickness interface elements

(Kim 1998; Zhang et al. 1999; Yang and Drumm 2000). Zhang et al. (1999) used two

fictitious springs perpendicular to each other to simulate the normal and tangential

behaviour of the nail-soil interface. Yang and Drumm (2000) adopted a type of surface

based frictional interaction to simulate the interaction between the nail and the

surrounding soil. The advantage of this method is that a large relative displacement

between soil and inclusion is allowed. But the actual behaviour of the nail-soil interface

is neither purely elastic nor frictional so that interface elements with more complex

behaviour need to be developed. The second method is using one dimensional structural

elements, such as beam and cable elements, directly imbedded in the soil mass to

simulate the inclusions (Lim 1996; Briaud and Lim 1997). The third method is to use a

thin layer of continuum elements to simulate the interface as reported by Tabrizi et al.

(1995). The advantage of this method is that many material properties can be selected

and there are fewer convergence problems than with the contact (interaction) method.

But only a finite relative displacement between the nail and the surrounding soil is

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allowed with the material properties remaining constant both before and after sliding of

the nail. In fact, the nature of the interaction between the nail and the surrounding soil is

a combination of cohesion and friction and the properties may not be the same before

and after sliding.

Table 2.1 – Basic assumptions of different soil nailing design approaches

Davis Method (Shen et al. 1981)

French Method (Schlosser 1983)

German Method (Stocker et al. 1979)

Kinematical Method (Juran et al. 1988)

Analysis Limit force equilibrium

Limit moment equilibrium

Limit force equilibrium

Limit force equilibrium

Global stability Global stability Global stability Local stability Input material properties

Soil parameters (c', φ'), limit nail forces, lateral friction

Soil parameters (c', φ'), limit nail forces, bending stiffness

Soil parameters (c', φ'), lateral friction

Soil parameters (c', φ'), limit nail forces, lateral friction

Nail forces Tension Tension, shear, bending

Tension Tension

Failure surface

Parabolic Circular, any input shape

Bilinear Log-spiral

Safety factors soil strength Fc, Fφ

1.5 1.5 1 (residual shear strength) 1

Pullout resistance Fp

1.5 1.5 1.5-2.0 2

Tension Yield stress Yield stress Yield stress Yield stress Bending Plastic moment Plastic moment Ground water No Yes No Yes Soil stratification No Yes No Yes

Loading Uniform surcharge Slope, any surcharge

Slope surcharge Slope surcharge

Structure geometry

Vertical facing Any input geometry Inclined facing, vertical facing

Inclined facing, vertical facing

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Figure 2.1 – Equipment for launched soil nails (After Myles and Bridle 1992) Figure 2.2 – Comparison of soil nailing, micro piles and soil dowelling (After Bruce and Jewell 1986)

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Figure 2.3 – Contrast of the construction sequence of reinforced earth and soil nailing (After Bruce and Jewell 1986) Figure 2.4 – Soil nailing mechanism (After Byrne et al. 1998)

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Figure 2.5 – Skin friction mobilization in pullout test (After Cartier and Gigan 1983) Figure 2.6 – Nails subject to shear and bending (After Mitchell 1987)

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Figure 2.7 – Davis design method (After Shen et al. 1981b) Figure 2.8 – Failure surfaces obtained by Davis design method and Finite Element analysis (After Shen et al. 1981b)

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Figure 2.9 – French design method (After Schlosser 1982) Figure 2.10 – Final yielding curve for inclusion in French method (After Schlosser 1982)

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Figure 2.11 – German design method: (a) Bilinear failure surface; (b) Acting forces and displacements; (c) Hodograph; and (d) Force polygon (After Gassler 1988) Figure 2.12 – Juran’s kinematical limit analysis design method: (a) Mechanism of failure; (b) State of stress in inclusion; and (c) Theoretical solution for infinitely long bar adopted for design purpose (After Juran et al. 1990)

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Figure 2.13 – Large model tests in the “Bodenvernagelung” project (After Gassler 1992) Figure 2.14 – Full scale test failure by breakage of inclusions in the “Clouterre” project (After Clouterre 1993)

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Figure 2.15 – Full scale test failure by reducing adherence length of inclusions in the “Clouterre” project (After Clouterre 1993) Figure 2.16 – Full scale test failure by excessive excavation in the “Clouterre” project (After Clouterre 1993)

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Figure 2.17 – Pullout box used by Tei (After Tei 1993) Figure 2.18 – Axial stress distributions along the nail obtained by Tei (After Tei 1993)

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Figure 2.19 – Pullout box used by Pradhan (After Junaideen et al. 2004) Figure 2.20 – Pullout box used by Chu (After Chu 2003)

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Figure 2.21 – Relationship between effect of reinforcement and inclination of inclusion (After Jewell 1987) Figure 2.22 – Failure surface of centrifuge model (After Tufenkjian and Vucetic 2000)

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Figure 2.23 – Reduced scale model test of a nailed soil slope (After Kitamura et al. 1988) Figure 2.24 – Reduced scale model test of a nailed wall (After Kim et al. 1996)

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Figure 2.25 – Mesh for (a) unreinforced slope and (b) soil nail slope developed by Yang and Drumm (2000)

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Chapter 3: Equipments and apparatus for laboratory pull-out tests

CHAPTER 3: EQUIPMENT AND APPARATUS FOR

PULL-OUT TESTS

3.1 PROBLEMS STUDIED BY LABORATORY PULL-OUT TESTS

As discussed in Chapter 2, the reinforced zone of a soil nail structure is separated into

an active zone and a passive zone during internal failure. In order to fully utilize the

strength of the steel bars, enough nail bond strength in the passive zone should be

ensured to avoid the nail being pulled out prematurely. Laboratory pull-out tests can be

used to study the pull-out resistance of a soil nail in the passive zone (or in the active

zone with the nail in an opposite pull-out direction).

Figure 3.1 shows a soil nailed slope. The deformation of the soil nailed slope (or the nail

itself) is a boundary value problem if no time effects (creep, consolidation, etc.) are

involved. To solve a boundary value problem, for example, the deformation and

stability analysis of a three-dimensional soil nailed slope, we need 3 types of equations:

(a) 3 stress equilibrium equations,

(b) 6 strain displacement compatibility equations, and

(c) 6 constitutive equations.

There are totally 15 equations for solving 15 unknowns (6 stresses, 6 strains and 3

displacements) by considering proper boundary conditions. The 6 constitutive equations

in fact describe the elementary stress-strain-strength behaviour of the soil (or interface).

As shown in Figure 3.1, to study the performance of a full length soil nail in a slope, we

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need a full-scale physical model (or a small-scale centrifuge model), since this is a

boundary value problem for this nail. On the other hand, if our purpose is to study the

influences of, for example, the installation procedures (hole drilling with stress release,

cement grouting) and overburden pressures on soil nail pull-out resistance, we need

only a representative short length of the nail as shown in Figure 3.1 (top left). This is

because that the full length of a nail in a slope at failure (the ultimate case for stability

analysis) is actually subjected to a tensile force. The interface shear stresses are in

opposite directions toward the slip surface as shown in Figure 3.1. The soil and stress

conditions along this full length nail are not uniform. Thus, if a full length nail with

nonuniform conditions is used for pull-out testing, the influences of parameters such as

overburden pressure and installation procedures cannot be clearly identified or separated

and therefore cannot be well studied.

However, if we consider a short length of the nail, the soil and stress conditions should

be approximately uniform along the longitudinal direction. We can then study the

influences of factors on the pull-out resistance of this short length of the nail in the soil.

The results obtained can be used as reference for a full length nail.

In the pull-out box study, a short length of the nail in a soil block is used (see the top

left in Figure 3.1). In such a test, the initial conditions of soil properties, stresses and

loading conditions in this short length in the longitudinal direction are all close to

uniform. This longitudinal section is considered to be an element, similar to the soil

specimen (element) used in a direct shear box or triaxial test. However, the stress and

loading conditions in the cross-section are not uniform, for example, due to hole drilling

and stress release, the radial stresses at the boundary of the hole are approximately zero.

The stresses and deformation in this cross-section are not uniform and are affected by

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the boundaries and loading. This is, in fact, what has been investigated in this study.

The nail with soil in the box is considered to be a boundary value problem with uniform

conditions in the axial direction (an element). The results obtained from the pull-out box

simulation can be applied to the full length of a soil nail in a slope subjected to the same

conditions.

3.2 NUMERICAL STUDY ON BOUNDARY EFFECT FOR DESIGN

OF THE BOX

With the discussion on the pull-out box simulation above, the design of the pull-out box

needs careful consideration of the boundary effects and an optimum selection of the size

(or dimension) of the box. The box should be large enough to ensure that the boundary

will not affect the pull-out test results. On the other hand, the box should not be too

large considering the weight of the box, the space required for the test and difficulties in

preparing a huge volume of soil sample.

In order to determine the distance between the nail centre and the boundary of the box, a

two-dimensional (2-D) plane strain finite element model was established using

ABAQUS code to obtain the stress redistributions in areas around the hole after drilling.

The dimensions of the model were 24m in height and 10m in width for simulating the

excavation of a hole with the diameter of 100mm (typical size in Hong Kong) at a level

of 20m beneath the ground surface as shown in Figure 3.2. A type of 4-node bilinear

quad reduced integration plane strain element was used in the model. The mesh was

very fine near the hole and relatively coarse when distant from the hole. The right and

left boundaries were located 5m away from the centre of the hole which was believed to

be sufficiently far that the influence caused by excavating the hole could be neglected.

The horizontal displacement of these two boundaries was restrained. The bottom of this

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model was located 4m from the centre of the hole and the vertical displacement was

restrained.

The simulation was accomplished in two analysis steps. Firstly, the initial stress

condition in the model was established by assigning an initial stress field and applying

gravity to the model to obtain an equilibrium state between them. The excavation of the

hole was simulated by removing the elements (here was the soil) at the location of the

hole using the “model change” command. Two constitutive models were used in the

simulation – an isotropic linear elastic model (as an elastic material) and an

elastic-perfectly plastic model using the Mohr-Coulomb failure criterion (as a

Mohr-Coulomb material).

Results from the ABAQUS finite element simulation are shown in Figures 3.3, 3.4 and

3.5. Figure 3.3 and Figure 3.4 show the vertical and horizontal stress contours induced

by the hole drilling procedure which were obtained by subtracting the stresses before

drilling from those after drilling. The stresses obtained were positive because the

compressive stresses were negative and decrease of compressive stresses (compressive

stresses are released during drilling) leads to increase of positive stresses. From the

contours in Figure 3.3 and Figure 3.4, it can be seen that the stresses induced by the

procedure of drilling the hole are negligible in areas which are over 300mm away from

the centre of the hole. Figure 3.5 shows the relationship between the vertical stress and

distance from the top circumference of the hole in the vertical direction before and after

drilling. For the Mohr-Coulomb material, the relative differences of the vertical stresses

before and after excavating the hole are 5.35% and 3.26 % at points with distances of

300mm and 400mm respectively from the top circumference of the hole in the vertical

direction. For the elastic material, the corresponding relative differences are 3.09% and

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1.87% as shown in Figure 3.5. In order to study the influence of the boundary effect,

another model with the same dimensions of the pull-out box was established and

simulated results are shown in Figure 3.6. The results showed that agreement between

the results from the two models are good.

From the above results, we can determine that the influence of drilling the hole is small

enough and can be neglected in areas which are over 300mm away from the center of

the hole. According to this conclusion and considering that there is a little larger

influence in the vertical direction due to the overburden pressure, the effective internal

dimensions of the box were determined as 0.6m wide, 0.83m high and 1.0m long. A

pull-out box with these dimensions was designed and two were manufactured to

accelerate the testing program. Considering the shortage of space in the laboratory, the

time and costs together with the accuracy of the test, the dimensions selected are

believed to be the optimum ones for our study purposes. But it should be pointed out

that the width of the box is only 6 times of the nail diameter so that the increase in

constrained dilatancy of the soil by boundaries of the box during pull-out may not be

neglectable. It should be careful to directly use the results from this project to field

conditions.

3.3 DESIGN AND CONSTRUCTION OF THE PULL-OUT BOX

3.3.1 Investigations to be conducted using the boxes

The pull-out boxes used for the study propose had to meet certain requirements. Before

testing, CDG soil is compacted into a box, normally to 95% of the maximum dry

density. A hole is then drilled after application of the vertical load. After drilling, a steel

bar is installed into the hole and kept centered. The hole is then grouted with cement

grout. After about 5 curing days until the strength of the cement grout has developed a

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value of at least 21MPa, the soil nail is ready to be pulled out with a hydraulic jack.

Following studies were carried out to investigate the influence of the following four

factors on the interface shear strength between the grouted nail and the surrounding soil:

(a) Overburden pressure (vertical stresses) of the soil

(b) Degree of saturation of the soil (dry or wet) and suction effects

(c) Influence of installation procedures (hole drilling, stress release, pressure grouting)

(d) Interface shear dilation

A box which can be used for investigating all of the above aspects has to fulfill some

fundamental requirements. Firstly the box should be strong enough to sustain the weight

of the soil and a higher applied overburden pressure with relative small deformation.

Secondly the box should be able to be sealed so that back water pressure can be applied

to the test soil to obtain a higher degree of saturation of the soil. A third requirement is

that enough access holes should be available for transducers cables and these holes must

be capable of being sealed when saturating the soil. All the above requirements were

allowed for during the design of the pull-out box.

3.3.2 Description of the pull-out box

The pull-out box was designed as shown in Figure 3.7 to Figure 3.10. The internal

dimensions of the box are 1.0m in length, 0.6m in width and 0.83m in height. A rubber

diaphragm is fixed under the bottom surface of the top cover of the box to form a fluid

chamber for application of overburden pressure to the top soil surface. A wooden board

is placed between the rubber diaphragm and the top soil surface in order to make the

soil deformation more uniform. The dimensions of the test soil are 1.0m in length, 0.6m

in width and 0.8m in height. The top 30mm thickness of space remaining above the soil

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surface in the box is used to house the rubber diaphragm and wooden board.

The box was constructed with five 8mm thick steel plates welded together. Square

section steel pipe was welded to the outside of the box to increase strength and reduce

bending deformation of the box. The top cover of the box was similarly constructed

with the 8mm thick steel plate and strengthened with steel pipes. A sheet of rubber

which was cut from a truck tyre inner tube was fixed under the top cover of the box by a

rigid steel frame using a quantity of short screws to form the fluid chamber for the

application of normal pressure. Two valves were connected to two holes on the top

cover of the box (see Figure 3.8 to Figure 3.10). One valve was for applying the

overburden pressure and included a pressure dial gauge for measuring the overburden

pressure; the other was for releasing air from the fluid chamber which may otherwise

lead to inaccurately measured volume changes in the test soil. This volume change

during the test was measured by an automatic volume change apparatus connected to

the pressure-supply line. The water pressure in the fluid chamber and the water pressure

source should be in equilibrium. Expansion of the soil sample would lead to an increase

in the water pressure in the fluid chamber and water would flow out of the fluid

chamber because of the nonzero hydraulic gradient. In contrast, water would flow into

the fluid chamber when the soil sample contracted. The volume of the water flowing

into and out of the fluid chamber was measured by the automatic volume change

apparatus and thus the volume change of the test soil was obtained. Four thicker (about

18mm) steel plates were welded to the outer surface of the box as a brim stiffener to

allow fixing of the top cover (Figure 3.8, 3.10 and 3.11). There were about 20 holes on

the brim stiffener and the cover of the box at the same locations. The top cover was

connected to the brim stiffener using M16 steel bolts during the tests. The brim stiffener

is a little bit lower than the top edge of the side plates of the pull-out box, to allow the

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rubber sheet to be squeezed tightly between the top cover and edges of the side plates as

tightening the bolts. This was able to prevent leakage of water from the pull-out box in

tests under submerged conditions (Figure 3.11). Four small 12mm thick steel plates

were welded to the strengthening beams of the front plate for connecting the drilling

machine during drilling the hole and the load reaction frame during the pull-out process

(Figure 3.8). Four thick rubber blocks were located under each foot of the box in order

to reduce vibration of the box during compaction and drilling. The rubber blocks could

also reduce disturbance (vibration and noises) caused by the test to the offices under or

surrounding the laboratory.

There are six 8mm diameter holes in the front and two in each side plate of the box as

accesses for wires of the pressure cells and tensiometers (or porewater pressure

transducers). The holes were sealed with watertight bolts (Figure 3.12) under

submerged conditions. The watertight bolt can be divided into two components and a

small rubber O-ring can be put between them. The two parts of the bolt are both hollow

and the transducer wire can pass through the hole along the bolt and be sealed by

tightening the two parts and squeezing the O-ring between them. On each side plate,

there are five holes for use in applying back pressure and for releasing air in the soil.

3.3.2.1 Extension cylindrical chamber covering the soil nail end

An extension cylindrical chamber of 250mm in length and 180mm in diameter (internal

space) was attached to the end of the box to cover the end of a soil nail as shown in

Figure 3.7, 3.8 and 3.10. This chamber was constructed using an 8mm thick steel

cylinder together with a flange welded on one end and an 8mm thick circular steel plate

fixed by bolts on the other end. This chamber was filled with the same type of soil

during the test. A hole was drilled through the box leading into this chamber. A soil nail

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was grouted in the drillhole and inside the chamber. The soil inside the extension

chamber was removed after the nail had hardened. Therefore the part of the nail within

the test box was maintained a constant length of 1.0m at all time during the pull-out

tests and no cavity would be left behind the end of the test nail. Thus the deformations

of the soil surrounding the hole would be more uniform and no stress concentration

would occur because no cavity was left behind the end of the nail. Therefore, the

stresses on the soil nail surface and in the surrounding soil would be more uniform than

has been the case in the boxes used by others (Chang et al. 1996, Franzén 1998, Lee et

al. 2001, Junaideen et al. 2004, Pradhan et al. 2003, Chu 2003). The extension chamber

was connected on the back plate of the box using the flanges on them with a deformable

rubber O-ring in between to prevent leakage.

A 140mm diameter hole was cut through the front plate with its centre 315mm from the

bottom of the box for drilling the hole and grouting the nail. On the back plate of the

box, there was an opening at the same level with a diameter of 180mm (the same as the

internal diameter of the extension chamber), which was used to make it possible to

extend the nail into the extension chamber.

3.3.2.2 A waterproof front cap to covering the soil nail head

A waterproof front cap was designed and constructed to be attached to the front plate of

the box to cover the soil nail head as shown in Figure 3.13 (the length unit in this figure

is cm). This waterproof front cap was used to prevent water leakage and enable back

pressure to be applied to the soil in the box in a water submerged condition. The

waterproof front cap is similar to a traditional triaxial cell. It is formed by a Perspex

cylinder which is restrained by a steel plate and a steel ring using tie rods. A steel rod is

used to connect the nail head and the bolt for pulling out. Two copper rings and

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deformable rubber O-rings are used at the location where the rod passes through the

steel plate to prevent leakage. The internal surfaces of the two copper rings are inclined

in order to make it possible to adjust the direction of the rod during pull-out. There are

several holes in the steel plate for water filling purpose and as accesses for strain gauge

wires. An extension square plate was designed to connect the waterproof front cap to

the box. To make connection, a 198×188×12mm steel plate was welded to the front

plate together with four M12 bolts on it. A rubber O-ring was used between the

waterproof front cap, square plate and the box to prevent leakage. There are four

elongated holes in the square plate which allow adjustment of the triaxial cell in the

vertical direction (Figure 3.13).

3.3.2.3 Application of back pressure for saturation of the soil

A permanent soil nail structure will serve in various weather conditions including

intense rainfalls. Even for temporary soil nail structures, it is also possible to meet

heavy rain during construction. Considering the possible variations of underground

water levels and weather conditions, the behaviour of a soil nail structure when the

degree of water saturation changes should be investigated even if drainage is provided

during construction of the structure. Thus pull-out tests should be carried out in soils

with higher degrees of saturation and nearly saturated soils. For a submerged test in

which the soil is close to saturation, the soil needs to be saturated before the nail is

pulled out. In order to obtain a higher degree of saturation, back water pressure needs to

be applied to the soil sample because of the large volume of the sample.

Preventing leakage of water is difficult during the application of back water pressure to

a soil nail pull-out box especially at the location of the nail head because the nail needs

to be pulled out through the nail head. Some authors have adopted steel wire reinforced

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flexible plastic tubes to cover the nail head in submerged soil nail laboratory pull-out

tests (Pradhan 2003; Chu 2003). In their method, however, back water pressure can not

be applied to the soil sample because the sealing could be damaged very easily so that it

is difficult to obtain a relatively high degree of saturation. In Chu’s tests, the obtained

highest degree of saturation was 86%. Pradhan stated that a nearly saturated condition

was obtained but the sealing was easily damaged. Enlightened by the mechanism of a

traditional triaxial apparatus, the specially designed waterproof front cap described in

the above paragraphs was used to cover the nail head in submerged pull-out tests. The

waterproof front cap could prevent leakage of water under a relative high (more than

100kPa) pressure and allowed the nail to be pulled out with very low friction between

the rubber O-ring and the steel rod for pull-out. The arrangement for applying back

pressure and saturating the soil is shown in Figure 3.14.

In the design of this pull-out box, the use of an extension chamber at back and a

waterproof cap in front of the box is most important. It becomes possible to keep the

stress and deformation of the soil more uniform during the course of pulling out and to

achieve a higher degree of saturation (even fully saturated) in the soil sample.

3.4 MEASURES FOR REDUCING SIDE FRICTION OF THE BOX

The internal surface of the box is not very smooth and the friction cannot be neglected.

Many options were tried to reduce the side friction of the box. We first planned to use a

type of PVC sheet attached to the internal surface of the box to reduce the friction. In

order to verify the effect of this method, a serious of direct shear tests on the interface of

this type of PVC sheet and the CDG soil were carried out. But the measured friction

angle of the interface was very large and even larger than the internal friction angle of

the pure soil. After the tests, many scratches were found on the PVC sheets, because the

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PVC sheet was so soft that the soil particles had penetrated resulting in a large friction

angle. Based on this experience, a stainless steel sheet with smooth surface was adopted

and direct shear tests were carried out on the interface of the stainless steel sheet and the

soil. The obtained friction angle was still a little larger.

Lubricating oil was then considered for spreading onto the stainless plate, but the oil

would pollute the soil which could not be reused. On the basis of this idea, a type of

flexible plastic sheet (film) was used to separate the soil from the oil. This approach

included a combination of the following three measures:

(a) A smooth stainless steel sheet was attached to the internal surface of the box.

(b) The stainless steel sheet was covered with a layer of lubricating oil.

(c) After this, a flexible plastic sheet (film) was placed on the stainless steel sheet with

the lubricating oil in between.

Direct shear tests were conducted to measure the friction angle of the interface between

the flexible plastic sheet (film) and the stainless steel sheet with the lubricating oil in

between. It was found that the friction angle was only 6.9o as shown in Figure 3.15. The

flexible plastic sheets (film) and the stainless steel sheets with the lubricating oil

between them placed on the surface inside the box are shown in Figure 3.16. The soil

was compacted in the box and in full contact with the flexible plastic film. The plastic

sheets can move more freely together with the soil in the vertical direction during

compaction of the soil and application of the overburden pressure to the test soil sample.

In order to study the influence of side friction on pressure transfer from the top soil

surface to the soil inside the box, a three-dimensional (3-D) finite element model was

established. Only one quarter of the test soil needed to be simulated because of the

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symmetry of the geometry and loading. The model was 0.32m in width, 0.52m in length

and 0.8m in height. The interface of the flexible plastic sheet (film) and the stainless

steel sheet with the lubricating oil between them was simulated by thin continuous

elements 0.02m thick and considered as a frictional material (φ=6.9°) to the front and

to the right side of the test soil. At the bottom of the model, the vertical displacement

was restrained to simulate the bottom support of the pull-out box. At the left and back

sides, the displacements normal to the surfaces were restrained to simulate the

symmetry of the model. At the front and right hand boundaries, the vertical

displacements and displacements normal to the surfaces were restrained. The applied

overburden pressure to the top surface of the test soil was 100kPa. The mesh and model

boundary conditions is shown in Figure 3.17.

The results are shown in Figures 3.18 and 3.19. Figure 3.18 gives the contours of

vertical stress, from which, we can see that the vertical stress drops slightly from top to

bottom. Figure 3.19 shows the relationship between vertical stress and distance from the

bottom of the model along the path shown in Figure 3.18. The vertical stress is 100kPa

at the top of the model and about 93kPa at the bottom of the model. These results

indicate that the influence of side friction was small after taking the above measures for

reducing side friction.

3.5 INSTRUMENTATION AND MEASUREMENTS

Comprehensive instrumentation was used for the laboratory pull-out tests. A description

of the transducers used and their measurements is as follows.

(a) Measurement of overburden pressure and vertical settlement of the top soil surface

In these laboratory pull-out tests, the overburden pressure was applied by a rubber

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diaphragm attached under the bottom surface of the top cover of the box through an

air-water interface from an air pressure source. A pressure dial gauge with a maximum

capacity of 350kPa was fixed to the top cover of the box to measure the applied

overburden pressure (Figure 3.14 and Figure 3.20). An automatic volume change

apparatus (Figure 3.14 and Figure 3.20), which is normally used in traditional triaxial

tests, was used to measure the volume change of the soil. The apparatus was attached to

the plastic tube which connected the top cover of the box and the air-water interface as

shown in Figure 3.14. The water in-flow (compression) or out-flow (expansion)

quantity was measured by the automatic volume change apparatus, which represents the

volume change of the soil mass in the box. The average settlement of the top soil

surface was calculated by dividing the measured water volume change by the area of the

top soil surface.

(b) Measurement of local soil pressures

Six strain gauge based earth pressure cells (TML Model KDA-PA/KD-A) with a

maximum capacity of 1MPa were embedded in the soil in each pull-out box to measure

and monitor earth pressures (Figure 3.20). Two earth pressure cells were placed 40mm

below the bottom of the soil nail and two cells 45mm above the top of the nail. The

remaining two pressure cells were in the middle between the soil nail and the top soil

surface. The locations of the earth pressure cells are shown in Figure 3.21. The

overburden pressure distributions in vertical direction and along the axis of the soil nail

could be obtained from the pressure cells at different time and stages during testing.

Variations in earth pressure in the course of testing were closely monitored particularly

before and after the drilling of soil nail installation holes.

(c) Measurement of soil suction and porewater pressure

- 73 -


The research required tests to be carried out in CDG soil with different degrees of

saturation, from unsaturated to saturated conditions. Miniature tensiometers (soil

moisture probe 2100F with an operational range of pressures from -100kPa to 100kPa)

and porewater pressure transducers (Druck PDCR-81 miniature pore pressure

transducers with a maximum capacity of 1.5MPa) were employed in the tests to

measure the suction and porewater pressure of the soil (Figure 3.20). As shown in

Figure 3.22, a total of four soil moisture probes (or porewater pressure transducers)

were installed in positions close to the nail surface (about 25mm away from the nail

surface). Soil moisture probes with vacuum dial gauges were used to measure the

suction of the soil compacted in the box for tests in partially saturated soil. For tests

under saturated conditions, both the soil moisture probes and miniature porewater

pressure transducers were used for measuring pore pressure in the soil. When using the

soil moisture probes to measure porewater pressure, the vacuum dial gauge of the probe

was replaced by a type of transducer which can measure both suction and positive pore

pressure (up to 100kPa).

(d) Measurement of axial strain in the soil nail

Four strain gauges were glued to the steel bar, at spacing of about 300mm with the first

strain gauge located 50mm away from the nail head to measure the axial strain of the

soil nail. The locations where the strain gauges were to be glued were polished to be

smooth and planar so that the strain gauges could fully contact the steel bar and work

properly. A water proof strain gauge was used to ensure that the cement grout would not

affect the accuracy of the measurement. Extra care was taken during installation of the

strain gauges to make sure that the strain gauges were parallel to the axis of the steel bar

so that accurate axial strain of the steel bar could be obtained. The measured axial strain

of the steel bar would be used to deduce the frictional shear force distribution on the

- 74 -


surface of the soil nail.

(e) Measurement of the pull-out displacement of the soil nail

Two LVDTs as shown in Figure 3.7 and Figure 3.20 were installed at the nail head to

measure the pull-out displacement. The average of the two values was used.

(f) Measurement of the pull-out force

The pull-out force was measured using a load cell located between the hydraulic jack

and the pull-out reaction frame as shown in Figure 3.7 and Figure 3.20.

Readings of all transducers were taken automatically by a datalogger connected to a

computer (Figure 3.20). The datalogger used in the tests is a CR10X datalogger with a

Multiplexer which can connect up to 32 transducers at the same time. There are two

terminals (High and Low) with each channel and the voltage difference between the two

terminals is used as output of the transducer response. This method can eliminate the

influence of the variation of the voltage supported by the datalogger. A computer

program needs to be developed to input transducer coefficients and determine which

channel is to be used by each transducer. The data were temporarily stored in the

datalogger and then automatically collected by the computer at a certain time interval

such as 30 minutes.

3.6 DRILLING MACHINE AND CEMENT GROUTING TOOLS

3.6.1 Drilling machine

The test soil was firstly compacted into the box under certain controlled conditions such

as initial dry density, initial water content, etc. The top soil surface was then loaded to a

constant vertical pressure. After about twenty four hours the deformation of the soil

- 75 -


sample had become negligible and a hole was then drilled by an electrical drilling

machine as shown in Figure 3.23.

In previous pull-out tests by Chu (2003) and the author’s own trial test made previous to

the carrying out of the main tests in this project, the drilling machine was just fixed to

the box with G clamps. Some dead loads were placed on the base of the drilling

machine to stabilize it, but violent vibration was observed and accurate drilling was

impossible. To solve this problem, the base of the drilling machine was firmly fixed to

the floor using expansive bolts. At the same time, the drilling arm was secured to a steel

plate fixed on the box with four long bolts (Figure 3.23). This significantly reduced the

vibration of the drilling machine during drilling and guaranteed the quality of the

drillhole.

The hole was drilled through the two access holes on the box into the extension

cylindrical chamber. In order to ensure that the hole would be successfully drilled into

the extension chamber and would not touch the back plate of the box, the hole had to be

coaxial with the two access holes. If the soil was compacted before setting up the

drilling machine, this would be very difficult to ensure. Thus the drilling bit and drilling

bars were connected to the drilling machine and the height and inclination of the drilling

arm were adjusted before compaction of the soil (Figure 3.24) to be coaxial with the two

access holes. The drilling machine was then fixed in position. Figure 3.25 shows a hole

under drilling.

3.6.2 Equipment for cement grouting without and with pressure

The grouting procedure was accomplished using a plastic pipe which was attached to

the steel bar. During grouting, a square steel plate supported the steel bar and blocked

- 76 -


the drillhole to prevent leakage of cement grout. There is one larger hole in the center

and two smaller holes above and below it on the square plate. The larger hole is to

accommodate the steel bar and the smaller holes are for the plastic pipes. Before

grouting, a longer plastic grouting pipe was attached to the steel bar and fixed on the

square plate together with the steel bar. The steel bar passed through the center larger

hole and the plastic pipe passed through the lower smaller hole. A shorter plastic pipe

was then inserted into the upper smaller hole as a grout-return pipe for air release and

for checking whether the drillhole was full or not. The gaps between the steel bar and

plastic pipes and the holes were sealed with plasticene. After this, the steel bar was

installed together with the grouting pipe to the end of the drillhole. The square plate was

then fixed on the pull-out box using steel bolts. The steel bar was supported by a steel

pipe and the load reaction frame to keep it centered in the hole.

At the beginning, gravity grouting were tried by simply raising the grouting pipe to a

level of about two meters above the drillhole for the cement grout to flow into the hole

under the gravity head. But this method was so slow that the cement grout solidified and

blocked the pipe before grouting was completed. A Perspex cylindrical container with a

piston was therefore designed to push the cement grout into the drillhole. A hole at the

bottom of the container enabled the grouting pipe to be connected to the cylinder. A

modified soil sample extruder was employed for grouting the drillhole. Grout was

pumped in by a piston fixed on the extruder when the grout container was driven up by

an electric motor of the extruder. Using this method, we can successfully complete the

grouting procedure. It was not possible however to apply any pressure to the cement

grout (Figure 3.26).

In order to accomplish the study of pressure grouting, there were several problems

- 77 -


needed to be solved. The first was the design of a grouting apparatus which could

sustain a high pressure and keep the pressure constant for a relatively long period. This

problem was solved by using a steel cylinder and air pressure to inject the cement grout.

The bottom and cover of the cylinder were fixed with steel bolts and were sealed with

rubber membranes to prevent leakage. There are two holes of the cylinder, one in the

top cover for connecting the pipe from the air pressure source and the other at the

bottom for connecting the grouting pipe. Another problem was the preventing of

leakage between the steel bar and grouting pipes and the square plate used for blocking

the drillhole. To solve this problem, a new square plate was machined and a kind of

hard plastic pipe was used which made it possible to use screw and O-ring to seal the

gaps. A third problem was how to measure the grouting pressure. The pressure of the air

source could be measured, but the pressure in the hole was smaller than that of the air

source because of the viscosity of the cement grout. Therefore the magnitude of the

grouting pressure in the drillhole was unknown. Another difficulty in measuring the

grouting pressure was that we could not measure the pressure of the cement grout

directly using a pressure dial gauge because the grout would block the gauge. This

problem was solved by using a type of oil to separate the cement grout from the

pressure dial gauge and the grouting pressure was transferred to the gauge by the oil.

Two pressure dial gauges were used — one lay in front of the hole and the other at

back of the hole — to take an average of the pressure (Figure 3.27).

Cement grout was poured into the cylinder first and then the top cover of the cylinder

was fixed. A constant air pressure was applied on the top surface of the cement grout.

The air pressure was kept constant for a relatively long period. Using the specially

designed cylinder, the cement gout was easily grouted under controlled pressures.

- 78 -


3.7 SETUP OF THE BOX FOR SOIL NAIL PULL-OUT TESTING

The experimental arrangement of the box with full instrumentation and loading devices

is shown in Figure 3.28. The pull-out was accomplished by a hydraulic jack and a load

reaction frame which was connected on the front plate of the box as shown in Figure

3.28. Two LVDTs and a load cell were used to measure the pull-out displacement and

pull-out force respectively. The LVDTs had to be carefully adjusted to be parallel to the

steel bolt for pulling out.

3.8 SUMMARY AND CONCLUSIONS

A new soil nail pull-out box has been designed and two were manufactured to study the

mechanisms of soil nail pull-out resistance and influencing factors. A series of tests

have been carried out using the boxes. The results obtained are reliable, accurate and

meaningful. Based on the presentation above, the innovative features and capabilities of

the boxes are summarized as follows:

(a) There was comprehensive instrumentation in the soil and the nail for measuring the

necessary parameters and meeting study requirements.

(b) An extension chamber was used to keep the test length and stress condition of the

soil nail constant during pull-out.

(c) A waterproof front cap was placed to cover the soil nail head and to seal the nail so

that back water pressure could be applied to accelerate the saturation of the soil in

submerged tests. Using this method, the influence of the degree of saturation on soil

nail pull-out resistance could be easily examined.

(d) The vertical stress was applied by a flexible rubber diaphragm fixed under the top

cover of the pull-out box. A fluid chamber formed by the rubber diaphragm and the

top cover was filled with de-aired water. The water in/out-flow was automatically

- 79 -


measured by an automatic volume change apparatus. In this way, the vertical stress

acting on the soil surface was more uniform and the volume change of the soil in

the box (or settlement) could be measured. This design made the box compact and

simple to operate.

(e) Specially designed apparatus was used to carry out pressure cement grouting so that

the influence of the applied grouting pressure could be studied.

(f) The overburden pressure was applied before installation of the soil nail in the test.

In this way, the stress release influence of the drilling of the hole on the soil nail

pull-out resistance could be investigated.

Based on the experience of using the two boxes and the results obtained, the following

conclusions may be drawn:

(a) The two pull-out boxes are simple, compact and capable of carrying out soil nail

pull-out tests under controlled conditions and meeting the study requirements.

(b) The instrumentation is reliable and meets the measurement requirements.

(c) The waterproof front cap was effective in sealing the nail head and could be

subjected to a high back water pressure. The whole design is effective for

increasing the degree of saturation of the soil to about 98%.

(d) The usage of the specially designed apparatus for cement pressure grouting is a

simple and effective way for conducting pressure grouting in the laboratory.

(e) The results from soil nail pull-out tests using the two new boxes were reliable,

accurate and meaningful. The results are presented and interpreted in the following

chapters.

- 80 -


Figure 3.1 – A soil-nailed slope and pull-out box simulation Figure 3.2 – Mesh for boundary effect study

10m

4m

20m

Soil nails

Uniform conditions in longitudinal direction

Slip surface

A soil-nailed slope Pull-out box simulation

Nonuniform conditions in cross-section

- 81 -


0.22

02m

Figure 3.3 – Vertical stresses induced by the hole drilling procedure

0.2796m

Figure 3.4 – Horizontal stresses induced by the hole drilling procedure

- 82 -


Elastic material

0.00.20.40.60.81.01.21.41.61.82.02.22.42.62.83.03.23.43.63.84.04.2

0 100 200 300 400Vertical stress (kPa)

Dist

ance

from

the

hole

surfa

ce (m

) __

After excavation

Before excavation

Mohr-Coulomb material

0.00.20.40.60.81.01.21.41.61.82.02.22.42.62.83.03.23.43.63.84.04.2


Dist

ance

from

the

hole

surfa

ce (m

) _

After excavation

Before excavation

(a) (b) Figure 3.5 – Relationship between the vertical stress and the distance from the top surface of the drillhole in vertical direction for (a) elastic material and (b) Mohr-Coulomb material

- 83 -


Figure 3.6 – Relationship between the vertical stress and the distance from the top surface of the drillhole in vertical direction for (a) box size model (b) large size model

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50


Dist

ance

from

the

hole

surfa

ce (m

) __

After excavation

Before excavation

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50


Dist

ance

from

the

hole

surfa

ce (m

) __

After excavation

Before excavation

(a) (b)

- 84 -


1 m

0.8

m

Wooden Plate

Rubber Diaphragm

Strengthening Beam

Pressure Gauge

CDG Soil Sample

Cement Grout

Rebar

Extension Chamber

Waterproof Front Cap

Guided Bar

Coupler

LVDT

Load Reaction Frame

Load Cell

Hydraulic Jack

Tensiometers

Earth Pressure Cells

Pressure Cells

Cement Grout

Rebar

0.6 m

0.8

m

Tensiometers

Figure 3.7 – Layout of transducers and pull-out equipment

- 85 -


Figure 3.8 – Design of the pull-out box – 3-D view Figure 3.9 – Back and front views of the pull-out box

Flange for Connecting the Extension Chamber

Rubber Diaphragm

Plate for Connecting the

g Machine and Reaction Frame

Pressure Gauge

Accesses for Transducer

wires Air Release

DrillinOpenings for Applying

Back Pressure Plate for Connecting the Triaxial Cell

Extension ChambeBrim

Stiffener r

- 86 -


Figure 3.10 – Side view of the pull-out box

Access holes for tensiometers and pore water pressure transducers

Strengthening beams

Brim Stiffener

Rubber diaphragm Brim stiffener

Edge of the side plate

Rubber gasket

Figure 3.11 – Cross-section of the top cover and the pull-out box

- 87 -


Figure 3.12 – Watertight bolt

40

40

17

1.6

5

Screw thread（M16×2）

Screw thread（M16×2）

Opening for applying water pressure

Access for transducer wires

Steel guide bar for pull-out

1.6

O-ring

Figure 3.13 – Waterproof front cap

- 88 -


Waterproof front cap

Water pressure dial gauge

Automatic volume change Apparatus

Box B

Figure 3.14 – Setup for saturating the soil

Failure envelop of the direct shear tests

y = 0.121x

0

5

10

15

20

25

30

35

40

0 50 100 150 200 250 300 350Normal stress (kPa)

Shea

r stre

ss (k

Pa) _

φ=6.9°

Figure 3.15 – Results of direct shear tests on the interface between the stainless steel sheet and the flexible plastic film with lubricating oil in between

- 89 -


Plastic and stainless steel sheets with oil in between

Figure 3.16 – Method for reducing side friction Figure 3.17 – Mesh and boundary conditions for investigating side friction of the box

- 90 -


Path for Figure 3.18

Figure 3.18 – Vertical stress contour for small side friction

Vertical stress vs. distance from the bottom of the model

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

0 20 40 60 80 100 120Vertical stress (kPa)

Dist

ance

from

the

botto

m (m

) __

Figure 3.19 – Vertical stress variations with distance from the bottom of the model along the path in Figure 3.17

- 91 -


LVDT

Soil moisture probe

Pressure dial gauges

Pressure cell

Load cell hydraulic jack

Data logger

Volume Change apparatus

Computer

Figure 3.20 – Transducers and datalogger used in the tests

- 92 -


2 1

34

6 5

400 300300 200

190

100

225

485

315

435

265

100

4540

Figure 3.21 – Locations of the earth pressure cells

400300 300

225

100

225

2525

Figure 3.22 – Locations of the soil moisture probes (or pore pressure transducers)

- 93 -


Drilling machine

Drilling arm

Figure 3.23 – Setup of the drilling machine

Drilling bit

Drilling bars

Figure 3.24 – Adjustment of drilling bit and drilling bars to ensure that they pass the centers of the two holes

- 94 -


Drilling bit

Figure 3.25 – Drilling the hole

Setup of the rebar

Pipe for grouting

Grout-return pipe

Grout container

Soil sample extruder

Figure 3.26 – Grouting without pressure (gravity head only)

- 95 -


Grout into drillhole

Pressure into cylinder

Pressure Supply

Grout-return pipe

Pressure Gauge

Box B Box A

Grouting cylinder

Figure 3.27 – Grouting with pressure

- 96 -


Box B

Waterproof Front Cap

Datalogger

Two LVDTs Figure 3.28 – Setup of the box with full instrumentation and loading devices for soil

nail pull-out under submerged condition

Load Cell

Hydraulic Jack

- 97 -

Chapter 4: Material properties and test procedures

CHAPTER 4: MATERIAL PROPERTIES AND TEST

PROCEDURES

4.1 INTRODUCTION

In Hong Kong, the completely decomposed granite (CDG) soil is one of the most

common soils (another is CDV – completely decomposed volcanic soil) forming fill and

cut slopes. Soil nailing has been the most popular method for stabilizing slopes in Hong

Kong since the late 1980’s. It is therefore important to study the behaviour of soil nail

structures and nail-soil interaction in a CDG soil. The soil used in this study was taken

from a site in Tai Wai, Hong Kong.

The properties of the CDG soil and cement grout, calibration of transducers and the test

procedures are described in this chapter.

4.2 MATERIAL PROPERTIES

4.2.1 Basic properties of the CDG soil

A series of basic property tests were carried out following the procedures as described in

BS 1377: 1990 to determine the basic properties of the CDG soil. The basic property

tests include particle size distribution, compaction tests, specific gravity, and liquid and

plastic limit tests.

The particle size distribution of the soil was determined by wet sieving and hydrometer

tests following the procedures in BS 1377-2 (1990) and GEO REPORT No.36 (Chen

1992). The result of the particle size distribution is shown in Figure 4.1. According to

- 98 -


British Standards (BS 5930, 1981), this CDG soil is composed of 9.33% gravel, 62.51%

sand, 24.97% silt and 3.19% clay. The mean grain size d50 was 0.25mm. The coefficient

of uniformity Cu and coefficient of curvature Cz of the soil are 38.3 and 1.6 respectively.

The plastic limit wp and liquid limit wl of the soil are 27.3% and 35.5% respectively.

According to the soil classification system in BS 5930 (1981), the soil can be classified

as a yellowish brown, very silty sand.

A standard compaction test using a 2.5kg rammer and a container of 1000cm3 was

carried out. The relationship between the dry density and moisture content of the soil is

shown in Figure 4.2. The maximum dry density ρdmax of the soil obtained is 1.668

Mg/m3 with an optimum moisture content mopt of 19%. The specific gravity Gs of the

soil is 2.645. The basic parameters of the soil are summarized in Table 4.1.

4.2.2 Determination of the shear strength of the soil

Conventional consolidated drained triaxial tests on recompacted saturated soil

specimens and double cell triaxial tests on recompacted unsaturated soil specimens had

been carried out. The double cell triaxial test was carried out using an inner cell and an

outer cell with the soil specimen within the inner cell. This apparatus makes it possible

to measure the volume change of partially saturated soil specimens by measuring the

volume change of water inside the inner cell. The recompacted soil specimens were

taken from the test soil sample which was compacted in the box to 95% of the

maximum dry density.

4.2.2.1 Conventional triaxial tests on saturated soil specimens

Both consolidated undrained (CU) and drained (CD) triaxial tests were carried out on

saturated recompacted soil specimens. The size of the specimen was 100mm in height

- 99 -


and 50mm in diameter. The initial degree of saturation of the soil specimens was 80%

and then they were saturated to the degree of saturation of 98% under a back pressure of

200kPa. This saturation condition was achieved by checking that the value of coefficient

B was larger than 0.95 (BS 1377, 1990). The consolidation process followed

immediately after the saturation stage to bring the specimen to the required effective

stress state. The consolidation was allowed to continue until there was no further

significant volume change, and until at least 95% of the excess pore pressure had been

dissipated. Five different consolidation pressures were used in the tests, which were

40kPa, 80kPa, 120kPa, 200kPa, and 300kPa. Finally axial compression was performed

under either undrained or drained conditions at axial displacement rates of 0.3mm/min

for the undrained tests and 0.1mm/min for drained tests until an axial strain of 20% was

achieved.

Figure 4.3 (a) shows the relationship between the deviator stress q and axial strain aε

under five different confining (consolidation) pressures for CU tests. No distinct peak

stress is observed in the curves and the maximum shear stresses are observed at the end

of the tests, that is, the stress-strain behaviour is strain hardening. Figure 4.3 (b) and (c)

show the relationship of porewater pressure u with axial strain aε and the effective

stress paths respectively. Figure 4.4 shows the relationship between s' (( 31 σσ ′−′ )/2) and

t (( 31 σσ ′+′ )/2) at axial strains of 15% and 20%. The calculated shear strengths (c' andφ′ )

are close to each other and the values at the axial strain of 20% were adopted.

In the CD tests, the sample was drained at bottom and porewater pressure was measured

on the top. Excess porewater pressures were observed at small strains and dissipate to

almost zero at the end of the tests. This indicated that the adopted strain rate was not

- 100 -


small enough and the tests were not real drained tests. During calculation of the shear

strength, effective stresses were used by subtracting a corrected porewater pressure from

the measured total stress. Figure 4.5 and 4.6 show the results of the ‘CD’ tests. There are

still no apparent peaks on the deviator stress-axial strain curves, but the deviator stresses

tend to be constant when the strain is larger than 15%. Shear strengths at axial strains of

15% and 20% were calculated and similar magnitudes were obtained.

4.2.2.2 Double cell triaxial tests on unsaturated soil specimens

The double cell triaxial system is shown in Figure 4.7 and 4.8. The apparatus is similar

to a conventional triaxial system except that two water pressure cells are used. The outer

water pressure cell has an internal diameter of 230mm, height of 425mm and wall

thickness of 8mm. The inner water pressure cell has an internal diameter of 90mm,

height of 235mm and wall thickness of 6mm.

For conventional triaxial tests, the volume change of unsaturated soil specimen cannot

be measured because of the presence of air in the voids. The volume change of the

water inside the water pressure cell cannot represent the volume change of the soil

specimen because the cell is deformable. In the double cell system, the water pressure in

the inner water pressure cell is equal to that in the outer water pressure cell so that the

deformation of the inner cell is negligible. Thus the volume change of water inside the

inner cell can be used as the volume change of the soil specimen (Yin 2003).

Figure 4.9 to figure 4.11 show the (a) Deviator stress vs. axial strain, (b) volume strain

vs. axial strain and (c) s vs. t for double cell triaxial tests on soil specimens with 38%,

50% and 75% degrees of saturation respectively. The stress-strain behaviour is normally

straining hardening under lower confining pressures. Under a higher confining pressure

- 101 -


(300kPa), slightly strain softening behaviour is observed for the stress-strain curves.

Compressive volume strain is considered to be positive. The volume strains are

observed to be compressive at the beginning of the tests and become dilative after

certain values of the axial strain, the larger the confining pressure, the longer the

compressive part of the volume strain-axial strain curves. Because the soil specimens

are unsaturated, the measured cohesions are larger than those in conventional triaxial

tests because of the suction effect. The measured shear strength parameters from both

conventional and double cell triaxial tests are summarized in Table 4.2.

4.2.3 Properties of the cement grout

The cement grout used for the tests had a water cement ratio of 0.42. Uniaxial

Compressive Strength (UCS) tests were carried out on both cylindrical and cubic

specimens to obtain the properties of the cement grout. The specimens were cured for 5

days which was the same as the curing time for the cement grout of the soil nail in the

pull-out tests. Totally four strain gauges were attached to the surface of the cylindrical

specimen with two of them parallel and the other two perpendicular to the axis of the

specimens. Photographs of failed specimens are shown in Figure 4.12. Results of the

tests on cylindrical specimens are shown in Figure 4.13. The average uniaxial

compressive strength cσ of the cement grout for cylindrical specimens is 32.09MPa.

The secant Young’s modulus and the corresponding Poisson’s ratio are

12.59GPa and 0.21 respectively. A total of four cubic specimens were also tested and

the measured average compressive strength is 32.2MPa. The density of the cement grout

was found to be 1.886Mg/m

50E nv

3. The properties of the cement grout are summarized in

Table 4.1.

- 102 -


4.3 CALIBRATION OF TRANSDUCERS

A CR10X data logger was used to collect data from the test. Before conducting the test,

all the transducers were calibrated over the entire operational ranges.

The pressure cells were calibrated using a GDS (Geotechnical digital system) pressure

controller and a sealed Perspex cylinder (Figure 4.14). Pressure cells were put into the

water in the cylinder and sealed with a watertight bolt. Then water pressures from

50KPa to 500kPa with an increment of 50KPa were applied to the water in the cylinder.

The responses (differential voltages) of the pressure cells under each pressure were

recorded and the relationship between the applied pressures and the responses of the

pressure cells was established. Therefore calibration factors for converting the responses

from the datalogger into pressures were obtained. The porewater pressure transducers

and transducers of the tensiometers were calibrated in the same way except that

negative pressure was applied for calibrating the transducers of the tensiometers. The

load cell was calibrated by a standard proving ring using the loading system of a triaxial

apparatus. The load cell was put between the loading plate of the triaxial apparatus and

the proving ring. Raise the loading plate, and then load was mobilized gradually on the

proving ring. The responses corresponding to each load were recorded and factors for

converting the responses into loads was obtained by plotting the load-response curve.

The LVDTs were calibrated using a micrometer to obtain the calibration factors

converting the datalogger responses into displacements. Figure 4.15 shows the

calibration results for one of the pressure cells, the tensiometers and the LVDTs, and the

load cell.

The strain gauges had no need of calibration but a bridge had to be developed for

connecting them to the CR10X datalogger. A full bridge was developed as shown in

- 103 -


Figure 4.16. R1, R2 and R3 are three 120Ω resistances and Rg is the strain gauge with

the same resistance. The H, L and AG are connected to the High, Low and Ground

terminals of the data logger’s differential channel separately. The Vx is connected to an

excitation voltage which provides the voltage for the strain gauge. It is important that

the gauge be wired as shown with the wire from H connected at the gauge, and that the

leads to the L and AG terminals be the same length, diameter and wire type. With this

configuration, changes in wire resistance due to temperature occur equally in both arms

of the bridge with negligible effect on the output from the bridge.

The result of the full bridge measurement is the measured bridge output in millivolts

divided by the excitation in volts, which means the result obtained directly from the data

logger is:

⎟⎟⎠

⎞⎜⎜⎝

⎛

+−

+⋅=⋅

21

2

3

10001000RR

RRR

RVV

g

g

in

out (4.1)

When strain is calculated the direct ratio of the voltages (volts per volts not

millivolts per volts) will be used:

21

2

3 RRR

RRR

VV

g

g

in

out

+−

+= (4.2)

When the gauge is strained it will change the resistance by , the equation for

the bridge output is:

gRΔ

21

2

3 RRR

RRRRR

VV

gg

gg

strainedin

out

+−

Δ++

Δ+=⎟⎟

⎠

⎞⎜⎜⎝

⎛ (4.3)

Subtracting the unstrained result from the strained result gives : rV

g

g

gg

gg

unstrainedin

out

strainedin

outr RR

RRRR

RRVV

VV

V+

−Δ++

Δ+=⎟⎟

⎠

⎞⎜⎜⎝

⎛−⎟⎟

⎠

⎞⎜⎜⎝

⎛=

33

(4.4)

- 104 -


Because the resistance of the strain gauge is also 120Ω, which is equal to the

resistance of R3, substituting Rg for R3:

gg

gr RR

RV

Δ+

Δ=

24 (4.5)

Solving for strain:

r

r

g

g

VV

RR

214−

=Δ

(4.6)

Strain is calculated by dividing the gauge factor from the above equation and the

units are converted to micro strain by multiplying by 106.

)21(10410 66

r

r

g

g

VGFV

RGFR

−⋅

=⋅

Δ=με (4.7)

4.4 SOIL PREPARATION AND TEST PROCEDURES

4.4.1 Soil preparation

Firstly, the CDG soil, with 95% of the maximum dry density, was compacted in the box

and the extension chamber in 9 layers (each with a maximum thickness of 100mm)

(1.668Mg/m3). Before compaction the soil was mixed to the moisture content

corresponding to a given degree of saturation. It was then manually compacted in the

box using a rammer whose weight was about 12kg. To achieve a relative compaction Cr

of 95%, the weight of each layer was calculated based on the moisture content w and

maximum dry density maxdρ of the soil. The required mass M of the soil sample for a

given volume V is determined by the following equation

[ VwCM rd )1(max += ]ρ (4.8)

A total of 6 earth pressure cells were embedded in the soil at three different levels

- 105 -


during the soil compaction. The installation of the pressure cells was divided into three

steps. The first step was to excavate a small pit in the compacted soil and put a small

amount of fine sand into the pit. The second step was to put a pressure cell (50mm in

diameter and 10mm in thickness) into the pit and check its level and whether it was in

full contact with the fine sand. The final step was to cover the pressure cell with the

same fine sand. Figure 4.17 shows the compaction of soil and installation of pressure

cells.

4.4.2 Preparation of soil specimens for triaxial tests

The recompacted soil specimens for triaxial tests were cut from soil samples taken from

the compacted test soil in the box using core cutters as shown in Figure 4.18. The core

cutter is 180mm long with an internal diameter of 50mm and produces a specimen

100mm in length and 50mm in diameter. After all the required soil had been compacted

into the box to 95% of the maximum dry density, four M16 bolts and two hollow square

steel sections were installed to form a reaction frame. The core cutter was then jacked

into the soil using a hydraulic jack against the reaction frame. The use of a hydraulic

jack instead of a hammer to push the core cutter into the soil was to reduce disturbance

to the soil sample. Finally the core cutter together with the soil sample was removed

from the soil and kept in a sealed plastic bag for further use in the triaxial test.

4.4.3 Application of vertical overburden pressure

After all the required soil had been compacted in the box, the top cover was connected

onto the box. Overburden pressure was then applied to the test soil by water pressure

through the rubber diaphragm under the top cover. De-aired water was continuously

filled in the fluid chamber formed by the rubber diaphragm and the top cover to

compensate for the deformation of the soil. The change in volume of the water was

- 106 -


measured using the automatic volume change apparatus. The volume change and the

variation of pressure cell readings were recorded using the CR10X data logger.

4.4.4 Hole drilling and cement grouting

About 24 hours after the application of overburden pressure, any further volume change

had become negligible, which indicated the soil was close to an equilibrium state under

the applied overburden pressure. A 100mm diameter hole was then drilled from the

front of the box, through the soil horizontally all the way into the extension chamber.

After this, a 40mm ribbed steel rebar with 4 strain gauges was placed in the center of the

hole. The hole was then grouted with cement grout with a water-cement ratio of 0.42.

Detailed description of arrangement of the drilling and grouting apparatus can be found

in Chapter 3.

4.4.5 Installation of tensiometers and/or porewater pressure transducers

In partially saturated soil, the tensiometers were installed after about 4 curing days (one

day before pull-out). A 2100F soil moisture probe shown in Figure 4.19 was used to

measure the soil suction in the case of partially saturated soil. Before being installed in

the soil, the probe was saturated with de-aerated water and the air which remained in the

vacuum dial gauge was sucked out under a negative pressure created by a GDS pressure

controller (Figure 4.20). Four 6mm holes, the same diameter (6mm) as the porous

ceramic tip, were drilled through the existing 8mm holes in both the side plates of the

pull-out box, extending to a location about 25mm away from the soil nail surface. The

plastic body tube of the probe was then fixed on the side of the box together with the

porous ceramic tip installed into the hole (Figure 4.19). The hole was finally filled up

with soil and sealed by watertight bolt (Figure 4.19). After about 24 hours, the soil

moisture probe and the surrounding soil reached equilibrium and the soil suction value

- 107 -


was read from the vacuum dial gauge. For tests in saturated soil, two tensiometers with

transducers and a porewater pressure transducer were installed before submerging the

soil to measure the porewater pressure. The same procedure was followed during

installation.

4.4.6 Saturation of the test CDG soil

For tests under submerged conditions, the soil was saturated before pull-out. About two

days after grouting, the cement grout had hardened and the steel plate for blocking the

hole during grouting was removed. Then the waterproof front cap (Figure 3.12) was

carefully fixed to the front plate of the box to cover the nail head. During installation,

the location of the waterproof front cap was carefully adjusted to ensure that the steel

guide bar could move smoothly during pull-out without any leakage. The soil in the

back extension chamber was then removed. Plastic pipes were connected to the holes on

the waterproof front cap, back extension chamber and left and right sides of the box.

After that a vacuum pump was used to suck out the air inside the soil from the upper

holes of the box and water was filled into the box from the lower holes at the same time.

After the waterproof front cap and the back chamber were full of water and not much air

was being sucked out, back pressure was incrementally applied to the soil to obtain a

higher degree of saturation. About two days later, no more water could be filled into the

soil under a given amount of back pressure and then the nail was ready to be pulled out.

The setup for saturating the CDG sample is shown in Figure 3.13.

4.4.7 Pull-out of the nail

After five curing days when the strength of the cement grout developed was about

32MPa, the nail was pulled out using a hydraulic jack against a steel reaction frame.

The reaction frame was secured on the box and its height could be adjusted to fit the

- 108 -


location of the nail head. A 1.2m long M16 bolt was connected to the nail head for

pull-out. Before pull-out the peak pull-out load was estimated based on previous

experience and evenly divided into five loading increments. The load was applied step

by step and was held for about one hour at each loading step. After the peak pull-out

resistance was achieved, the nail was continuously pulled out under displacement

control using rates of 1mm/min or about 0.3mm/min for tests in partially saturated and

saturated soil respectively. For tests in partially saturated soil, the displacement at the

end of pull-out was 200mm. However, for tests in saturated soil, a displacement of only

100mm was achieved because of the difficulties in maintaining alignment of the

pull-out force with the nail and preventing leakage of water from the waterproof cap.

Time constraint was another reason because the displacement rate for pull-out in the

submerged tests was very slow and the time needed would be too long if proposed to

pull out of 200mm. The setup of the box with full instrumentation and loading devices

for soil nail pull-out under submerged condition is given in Figure 3.23. Figure 4.21

shows a soil nail being pulled out under a dry soil condition.

The pull-out of the nail was conducted by load control before peak pull-out resistance

and by displacement control after peak. The loading speed was slow and a relatively

long time interval of one hour was maintained between two loading steps. The

displacement rate after peak pull-out resistance was small, especially for saturated tests.

Therefore the pull-out tests could be considered to have been carried out under

approximately drained conditions.

- 109 -


Table 4.1 – Properties of the CDG soil and cement grout

Properties of the completely decomposed granite soil Specific Gravity (Gs) 2.645 Maximum dry density (ρdmax) Mg/m3 1.668 Optimum moisture content % 19 Plastic limit (wp) % 27.3 Liquid limit (wl) % 35.5 Gravel % 9.33 Sand % 62.51 Silt % 24.97 Clay % 3.19 Coefficient of uniformity Cu 38.3 Coefficient of curvature Cz 1.6 Properties of the cement grout Density (ρd) Mg/m3 1.886 Uniaxial compressive strength (σc) MPa 32.09 Secant Young's modulus (E50) GPa 12.59 Poisson's ratio (υ) 0.21

Table 4.2 – Shear strength parameters of the CDG soil

Degree of Saturation Shear strength

Type of Test (Sr)

Cohesion (c or c')

Friction Angle (φ or φ')

Sample Condition

38% c=36.6kPa φ =35.9° Recompacted

50% c=59.5kPa φ =30.4° Recompacted Unsaturated Consolidated

Drained (CD) 75% c=26.8kPa φ =33.8° Recompacted

Consolidated Drained

(CD) 98% c’=9.4kPa φ'=33.9° Recompacted

Saturated Consolidated Undrained

(CU) 98% c’=11.4kPa φ'=32.6° Recompacted

- 110 -


Particle Size Distribution

0

10

20

30

40

50

60

70

80

90

100

0.001 0.01 0.1 1 10 100

Particle Size (mm)

Perc

enta

ge F

iner

0.002 0.06 2 60

Fine Medium Coarse Fine Medium Coarse Fine Medium CoarseClay

Silt Sand Gravel Cob

ble

Figure 4.1 – Particle size distribution of the CDG soil

Relationship between dry density and moisture content

1.50

1.521.54

1.561.58

1.60

1.621.64

1.661.68

1.70

5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30Moisture content (%)

Dry

den

sity

(Mg/

m3 )

2.5 kg rammer

ρ d max=1.668Mg/m3

w opt =19%

Figure 4.2 – Relationship between dry density and moisture content

- 111 -


0

100

200

300

400

500

600

0 0.05 0.1 0.15 0.2Axial strain, εa

Dev

iato

r Stre

ss (k

Pa) 40kPa

80kPa

120kPa

200kPa

300kPa

(a)

-50

0

50

100

150

200

0 0.05 0.1 0.15 0.2Axial strain, εa

Pore

wat

er P

ress

ure,

u (k

Pa) _ 40kPa

80kPa

120kPa

200kPa

300kPa

(b)

0

200

400

600

800

0 100 200 300 400 500 600 700 800 900 1000Mean stress p' (kPa)

Dev

iato

r stre

ss q

' (kP

a) 40kPa

80kPa

120kPa

200kPa

300kPa

(c) Figure 4.3 – (a) Deviator stress vs. axial strain (b) pore water pressure vs. axial strain and (c) effective stress paths for conventional saturated CU triaxial tests

- 112 -


y = 0.5429x + 12.16

0

100

200

300

400

0 100 200 300 400 500 600 700 800 900 1000

t' (kPa)

s' (k

Pa)

t'=(σ'1+σ'3)/2

s'=(σ1-σ3)/2

c'=14.5kPa

φ '=32.9°

(a)

y = 0.5388x + 9.6163

0

100

200

300

400

0 100 200 300 400 500 600 700 800 900 1000t' (kPa)

s' (k

Pa) t'=(σ'1+σ'3)/2

s'=(σ1-σ3)/2

c'=11.4kPa

φ '=32.6°

(b) Figure 4.4 – Relationship between s' and t' for conventional saturated CU triaxial tests at axial strain of (a) 15% and (b) 20%

- 113 -


0100200300400500600700800900

0 0.05 0.1 0.15 0.2Axial strain, εa

Dev

iato

r Stre

ss (k

Pa) 40kPa

80kPa

120kPa

200kPa

300kPa

(a)

-50

0

50

100

150

200

0 0.05 0.1 0.15 0.2Axial strain, εa

Pore

wat

er P

ress

ure,

u (k

Pa) _

40kPa

80kPa

120kPa

200kPa

300kPa

(b)

0

200

400

600

800

1000

0 100 200 300 400 500 600 700 800 900 1000Mean stress p' (kPa)

Dev

iato

r Stre

ss q

' (kP

a) 40kPa

80kPa

120kPa

200kPa

300kPa

(c) Figure 4.5 – (a) Deviator stress vs. axial strain (b) pore water pressure vs. axial strain and (c) effective stress paths for conventional saturated ‘CD’ triaxial tests (partially drained and correction made on the excess pore water pressure)

- 114 -


y = 0.5829x

0

100

200

300

400

500

0 100 200 300 400 500 600 700 800 900 1000t' (kPa)

s' (k

Pa)

t'=(σ'1+σ'3)/2

s'=(σ1-σ3)/2

c'=0

φ '=35.7°

(a)

y = 0.5575x + 7.8021

0

100

200

300

400

500

0 100 200 300 400 500 600 700 800 900 1000t' (kPa)

s

Figure 4.6 – Relationship between s' and t' for conventional saturated ‘CD’ triaxial tests at axial strain of (a) 15% and (b) 20% (partially drained and correction made on the excess pore water pressure)

' (kP

a)

t'=(σ'1+σ'3)/2

s'=(σ1-σ3)/2

c'=9.4kPa

φ '=33.9°

(b)

- 115 -


Figure 4.7 – Schematic diagram of the double cell triaxial system Figure 4.8 – The double cell triaxial system (After Yin 2003)

- 116 -


0

200

400

600

800

1000

1200

0 5 10 15 20 25Axial Strain, εa (%)

Dev

iato

r Stre

ss (k

Pa)

80kPa

120kPa

200kPa

300kPa

(a)

-5

-4

-3

-2

-1

0

1

2

0 5 10 15 20 25

Axial Strain, εa (%)Vol

ume

Stra

in, ε

v (%

) _ 80kPa

120kPa

200kPa

300kPa

(b)

y = 0.5863x + 29.658

0

100

200

300

400

500

600

0 200 400 600 800 1000t (kPa)

s (kP

a)

t=(σ1+σ3)/2s=(σ1-σ3)/2

c=36.6kPaφ=35.9°

(c) Figure 4.9 – (a) Deviator stress vs. axial strain (b) volume strain vs. axial strain and (c) s vs. t for double cell triaxial tests on soil specimens at 38% degree of saturation

- 117 -


0100200300400500600700800900

0 5 10 15 20 25Axial Strain, εa (%)

Dev

iato

r Stre

ss (k

Pa)

80kPa120kPa200kPa300kPa

(a) -6

-5

-4

-3

-2

-1

0

1

2

3

0 5 10 15 20 25

Axial Strain, εa (%)

Vol

ume

Stra

in, ε

v (%

)


(b)

y = 0.5061x + 51.298

0

100

200

300

400

500

0 200 400 600 800t (kPa)

s (kP

a) t=(σ1+σ3)/2s=(σ1-σ3)/2

c=59.5kPaφ=30.4°

(c) Figure 4.10 – (a) Deviator stress vs. axial strain (b) volume strain vs. axial strain and (c) s vs. t for double cell triaxial tests on soil specimens at 50% degree of saturation

- 118 -


0100200300400500600700800900

0 5 10 15 20 25Axial Strain,εa (%)

Dev

iato

r Stre

ss (k

Pa)

80kPa120kPa200kPa300lPa

(a)

-6

-5

-4

-3

-2

-1

0

1

0 5 10 15 20 25

Axial Strain, εa (%)

Vol

ume

Stra

in, ε

v (%

)


(b)

y = 0.5563x + 22.271

0

100

200

300

400

500

0 200 400 600 800t (kPa)

s (kP

a) t=(σ1+σ3)/2s=(σ1-σ3)/2

c=26.8kPaφ=33.8°

(c) Figure 4.11 – (a) Deviator stress vs. axial strain (b) volume strain vs. axial strain and (c) s' vs. t for double cell triaxial tests on soil specimens at 75% degree of saturation

- 119 -


Figure 4.12 – Failed cement grout specimens of Uniaxial Compressive Strength (UCS) tests

- 120 -


Uniaxial Compressive Strength for Sample A

0

5

10

15

20

25

30

35

-1.E-03 0.E+00 1.E-03 2.E-03 3.E-03 4.E-03 5.E-03

)

Axi

al S

tress

(MPa

Uniaxial Compressive Strength for Sample B

0

5

10

15

20

25

30

35

-2.E-03 -1.E-03 0.E+00 1.E-03 2.E-03 3.E-03 4.E-03 5.E-03 6.E-03

Axi

al S

tress

(MPa

)

σa=33.0MPa

σa/2=16.5MPa

Δεa=1.30e-03

α

Δεl=-2.80e-04

Axial StrainLateral Strain

Figure 4.13 – Results of Uniaxial Compressive Strength (UCS) tests for the cement grout (cylindrical specimen)

σa=31.2MPa

σa/2=15.6MPa

Δεa=1.25e-03Δεl=-2.50e-04

α

Axial StrainLateral Strain

- 121 -


Watertight bolt

GDS pressure controller

Figure 4.14 – Apparatus for calibrating earth pressure cells, pore water pressure transducers and transducers for the soil moisture probes

Pressure cell1(EBJ04026)

y = 441.45x - 1197.5

0100200300400500600

0 2 4 6Dataloger reading (mv)

Pres

sure

(kPa

)

Tensiometer

y = -2.1037x + 2.8473

-120-100-80-60-40-20

00 20 40

Dataloger reading (mv)

Pres

sure

(kPa

)

60

LVDT1(m352407)

y = 13.51x + 1.2052

020406080

100120

0 2 4 6 8Dataloger Reading (mv)

Dis

plac

emen

t (m

m)_

Load cell

y = -12324x - 2327.5

0

1

2

3

4

5

-0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0Dataloger reading (mv)

Load

(kN

)

Figure 4.15 – Calibration results for the transducers

- 122 -


Figure 4.16 – Full bridge connection for strain gauge

H

L

AG

Vx

R1

R2

R3

Rg

Pressure cells

Access holes

Rammer

Leveler

Fine sands

Figure 4.17 – Soil compaction and pressure cell installation

- 123 -


Reaction frame

Hydraulic jack

Core cutters

Remove the core cutter out of the soil

A core cutter with a sample

Figure 4.18 – Taking soil samples for triaxial tests

- 124 -


Vacuum Dial Gauge

Porous Ceramic tip

Plastic Body Tube

Watertight bolt

Plastic pipe

Body tube fixed on the box

Figure 4.19 – Tensiometer (Soil Moisture Probe) used in the test Figure 4.20 –De-airing for a tensiometer

- 125 -


Figure 4.21 –A soil nail being pulled out under a dry soil condition

- 126 -


CHAPTER 5: INFLUENCE OF OVERBURDEN PRESSURE ON

SOIL NAIL PULL-OUT BEHAVIOUR AND RESISTANCE

5.1 INTRODUCTION

There are a number of factors that govern the shear strength between a soil nail and the

surrounding soil, such as the stress condition, soil properties, groundwater condition,

soil nail surface condition and etc. As mentioned in Chapter 2, the stress conditions that

will influence the pull-out resistance of a soil nail are normally referred to the normal

stresses acting on the nail surface. Current views on how the overburden pressure

influences the normal stress acting on the nail surface are diversified. Some researchers

believe that the normal stress is related to the soil overburden pressure corresponding to

the depth of soil, such as Jewell (1990). Some other researchers hold the view that the

normal stress is independent of soil depth, such as Schlosser and Guilloux (1981) and

Cartier and Gigan (1983).

In the design approach commonly adopted in Hong Kong, the pull-out shear resistance

of a soil nail is assumed to be directly proportional to the vertical stress. The vertical

stress that acts on the soil nail is calculated as the overburden stress at the mid depth of

the nail in the resistant zone (GEO 2005). However, for cement grouted nails that are

commonly used in Hong Kong, when a drillhole is formed for the soil nail installation,

the soil stresses on the surface of the hole are released and the soil above the hole is

supported by soil arching developed across the hole. As just a certain amount (not all) of

the stresses may be restored after hardening of the grout, the current design method may

not be able to reflect the actual site condition. Quality field test and laboratory test

- 127 -

Chapter 5: Influence of overburden pressure on pull-out resistance

results in this area are limited, especially for soil nails in the completely decomposed

granite (CDG) soils in Hong Kong. The CDG soils are one of the most common types

of soils in Hong Kong, covering more than 70% of the territories of Hong Kong.

In order to study the influence of overburden pressure on the soil nail pull-out shear

resistance, a number of pull-out tests in the CDG soil with various degrees of saturation

under different overburden pressures have been carried out using the two boxes

introduced in Chapter 3. The results of the tests in the soil at 38%, 50%, 75% and 98%

(submerged) degrees of saturation and under applied overburden pressures of 40kPa,

80kPa, 120kPa, 200kPa and 300kPa are presented in this chapter. The stress release,

development of normal stress, pull-out stress-displacement behaviour and influence of

applied overburden pressure on the pull-out resistance of soil nails will be presented and

discussed in the following sections.

5.2 STRESS VARIATIONS DURING DRILLING AND GROUTING

5.2.1 Stress release during drilling

The responses of earth pressure cells and the automatic volume change apparatus during

application of overburden pressure for all the tests are of similar pattern. Figure 5.1

shows the (a) total earth pressure and (b) vertical displacement versus time during

application of overburden pressure for the test under overburden pressure of 200kPa and

at initial degree of saturation (Sr) of 38%. The overburden pressure reached the target in

a short time and kept constant before drilling the hole. The change in volume of the soil

started to be measured as soon as the overburden pressure had reached its target. The

vertical displacement was calculated from the measured volume change of the soil.

From Figure 5.1, it can be observed that the vertical displacement increased fast firstly

and then slowed down with time elapsed. The increasing rate of the vertical

- 128 -


displacement became substantially small about ten hours (600min) after application of

the overburden pressure, indicating that the soil was close to an equilibrium state under

the applied overburden pressure. This state was held up to 24 hours to further improve

the equilibrium of the soil.

The results of the tests with applied overburden pressure of 200kPa and at degrees of

saturation of 38% and 75% are used to illustrate the variations of earth pressure during

drilling and grouting (Figure 5.2). The variations of earth pressure for tests in dryer and

wetter soil are both included. The locations of the earth pressure cells are shown in

Figure 3.20. From Figure 5.2 it can be seen that, before drilling, the earth pressures in

the soil as measured by the pressure cells generally agreed with the applied overburden

pressure of 200kPa (the slight difference was due to the self-weight of soil above the

earth pressure cells). The changes in earth pressures measured by earth pressure cells 5

and 6 were not significant comparing with the other four pressure cells. The earth

pressures measured by earth pressure cells 1 to 4 substantially decreased soon after the

drilling bit had passed them. Stress redistribution in the soil surrounding the hole was

apparent. In Figure 5.2 (a), the earth pressures measured by pressure cell 2 and pressure

cell 4 dropped to about 70kPa and those measured by pressure cell 1 and pressure cell 3

dropped to about 50kPa and 60kPa respectively. In Figure 5.2 (b), the earth pressures

measured by pressure cell 1 and pressure cell 2 dropped to about 70kPa and those

measured by pressure cell 3 and pressure cell 4 dropped to about 20kPa. Residual

pressures were recorded because the pressure cells were about 40mm away from the

surface of the hole where stresses remained locked in the soil due to the arching effect.

It can be observed that, for the test in soil at 75% degree of saturation, the residual earth

pressures measured by pressure cells 3 and 4 were smaller than those of pressure cells 1

and 2. This phenomenon was observed in most of the tests and was probably because

- 129 -


that pressure cells 3 and 4 were, located above the hole, affected by gravity. For the test

in soil at 38% degree of saturation, the residual earth pressure measured by earth

pressure cell 1 was smaller probably due to slight misalignment of drilling.

5.2.2 Variation of earth pressure during and after grouting

Upon injection of grout, the earth pressures were observed to increase slightly in some

of the tests but keep unchanged in the others. In the tests presented here, changes in the

earth pressures were not notable (Figure 5.2). Immediately after grouting, the earth

pressures started to decrease in almost all the tests probably due to the softening of the

soil by water of the cement grout. In Figure 5.2 (a), a remarkable decrease of more than

20kPa was observed immediately after grouting. A small decrease of about 5kPa was

observed in Figure 5.2 (b) after grouting. The decrease of earth pressures was normally

more significant in tests under higher overburden pressure. In the tests under higher

overburden pressure, the residual earth pressures after drilling were also higher so that

the decrease in earth pressures was more significant. The variation of earth pressures

before and after grouting was also different in tests with soil at different degrees of

saturation, which will be discussed in next chapter. One or two days after grouting, the

cement grout hardened and the earth pressures recovered a little bit perhaps owing to

stress redistribution. The recovered earth pressure was negligible in Figure 5.2 (a). In

Figure 5.2 (b), the earth pressures measured by pressure cells 1 and 2 recovered to the

values which were about 10kPa higher than those before grouting, and earth pressures

measured by pressure cells 3 and 4 recovered to the same value as those before grouting.

For all the tests, the recovered stresses are neglectable compared with the applied

overburden pressure.

The pull-out resistance of a soil nail is dependent on the local stress state of soil around

- 130 -


the drillhole at the time of pull-out. From the results, it is observed that the stresses in

the soil around the drillhole were largely released after drilling and the recovered

stresses after the grout had hardened was very small in comparison with the applied

overburden pressure. Thus in design of soil nailing system, the normal stress exerted on

the soil nail surface may not be taken as the weight of the soil above the soil nail as a

matter of course.

5.3 DEVELOPMENT OF EARTH PRESSURE DURING PULL-OUT

The average pull-out shear stress was calculated by dividing the pull-out force by the

surface area of the nail. The average earth pressure was calculated by averaging the

earth pressures measured by earth pressure cells 1 to 4. Figure 5.3 shows the typical

variations of average pull-out shear stress and average earth pressure against (a) time

and (b) displacement during a pull-out test in the soil at degree of saturation of 75% and

under an applied overburden pressure of 300kPa. The soil nail was pulled out by a

hydraulic jack using load control method. The loading increment was calculated by

dividing the peak load estimated from current design practice by a factor of 5. Because

the estimated peak load could not be perfectly equal to the actual peak pull-out load, the

amount of loading steps varied from test to test. There are four loading steps in Figure

5.3 (a). At each load increment the nail was held for about one hour. After the peak

stress was mobilized, the nail was pulled out continuously at a rate of 1.0mm per

minute.

From Figure 5.3, it can be seen that the average earth pressure increased simultaneously

with increase in the pull-out shear stress. No change in the earth pressure was observed

during the time interval when the applied load was held constant between two loading

increments. After the pull-out shear stress reached the peak, it started to decrease and

- 131 -


the same phenomenon happened to the earth pressure. The increase in earth pressure

was due to constrained dilatancy of the soil. The completely decomposed granite (CDG)

soil used in these tests is basically granular and was compacted to a very dense state. It

is therefore expected to dilate under shearing. The soil around the nail was under the

constraint imposed by the nail and the surrounding soil. This caused an increase in earth

pressure as measured by pressure cells during pull-out. It is apparent that this increased

normal stress due to the effect of confined dilatancy contributes significantly to the

pull-out shear resistance for drill-and-grout nails in dense granular soils.

Figure 5.4 shows the changes of average earth pressure measured by earth pressure cells

1 to 4 at different stages of testing for all the tests with soil at 38% degree of saturation.

The variations of average earth pressure during the course of testing, which are similar

to what has been discussed in the above paragraphs, are clearly shown in this figure.

After application of overburden pressures, the average earth pressures measured by

pressure cells 1 to 4 were approximate to the applied overburden pressures. The average

earth pressures substantially decreased to residual values after drilling and did not

recover much after installation of the soil nail. At the peak pull-out resistance, the

average earth pressures increased to certain values due to the constrained dilation of the

soil. The increased average earth pressures were scattered and not directly related to the

overburden pressure.

It needs to point out that the pressure cells above and below the nail did not respond

consistently. The drillhole might not be strictly in the same alignment with the steel bar

of the soil nail. The pull-out load was therefore not perfectly parallel to the soil nail

surface so that the pressure cells did not respond consistently. For example, if the

pull-out load slightly inclined upward, the earth pressures measured by earth pressure

- 132 -


cells 2 and 3 should be larger than those measured by earth pressure cells 1 and 4.

Therefore the average value of readings for the four earth pressure cells (P-cells 1 to 4)

was used in the analysis.

5.4 PULL-OUT SHEAR STRESS-DISPLACEMENT BEHAVIOUR

Figure 5.5 to Figure 5.8 show the relationship between average pull-out shear stress and

pull-out displacement for tests with soil at degrees of saturation of 38%, 50%, 75% and

98% (submerged tests) respectively. Results of the tests under overburden pressures of

40kPa and 80kPa, in soil at 50% degree of saturation and under overburden pressure of

80kPa, in soil at 38% degree of saturation were not included because those tests were

believed to have failed due to grouting problem. In general, the pull-out shear stress

firstly develops rapidly at small initial pull-out displacements, and then it continues to

develop as the displacement increases but at a decreasing rate as it approaches the peak

shear stress. After the peak shear strength, the pull-out shear stress is observed to

decrease gradually towards its residual value at the displacement of 200mm. The

displacement and pull-out resistance of the tests do not agree with the results presented

by Pradhan et al. (2003) and Chu (2003). The discrepancy may be due to the difference

in stress states of soil around the test nail between theirs and the present study. In the

tests reported by Pradhan et al. (2003) and Chu (2003), the pressure was applied after

soil nail installation but in the tests in this project, the pressure was applied before soil

nail installation and the stress in soil on the surface of the drillhole was released during

drilling. As a result, the normal stress on the soil nail was less at the beginning of

pull-out. The pull-out stress was generally mobilized by constrained dilatancy of the soil

and thus the displacement required to mobilize the peak shear strength was larger

compared to their results. The cement grout would contract during hardening and thus

the applied overburden pressure after installation of the soil nail would improve the

- 133 -


bond between the nail surface and the surrounding soil. This should be another reason

for the smaller displacements at the peak shear strengths occurred in their tests.

The pattern of the pull-out shear stress-displacement curves is different for tests in soil

at different degrees of saturation. The displacements required to mobilize the peak

pull-out shear strengths decrease with the increase in degree of saturation of the soil.

The post-peak decrease of pull-out shear stress becomes more apparent with the

increase in degree of saturation of the soil. The reason for these phenomena will be

discussed in next chapter concerning the influence of degree of saturation of the soil on

the soil nail pull-out resistance and behaviour.

5.5 INFLUENCE OF OVERBURDEN PRESSURE ON PULL-OUT

SHEAR RESISTANCE

5.5.1 Peak pull-out shear resistance

As shown in Figure 5.5 to Figure 5.8, four series of tests in soil at different degrees of

saturation had been carried out to study the influence of overburden pressure on the

pull-out shear resistance. For soil used in the first series of tests, the degree of saturation

was 38% and the measured average soil suction was -87kPa. For soil used in the second

series of tests, the degree of saturation was 50% and the measured soil suction was

-68kPa. For soil used in the third series of tests, the degree of saturation was 75% and

measured average soil suction was -6kPa. The suction of the soil generally kept constant

during testing. It was measured for reference only and was not used for analysis. In the

analysis of pull-out tests results, total stress was used. For soil used in the fourth series

of tests, the soil was submerged by water under back pressure and an average degree of

saturation of 98% was achieved. Each series of tests was subjected to applied

overburden pressures of 40kPa, 80kPa, 120kPa, 200kPa and 300kPa respectively. The

- 134 -


peak shear resistances under different applied overburden pressures in these four series

of tests are shown in Figure 5.9 to Figure 5.12. It can be seen that, for tests in soil at

degree of saturation of 38%, the peak pull-out shear resistances were quite constant

under different overburden pressures. For tests under submerged condition, the peak

pull-out resistances were also generally constant. However, for tests in soil at degrees of

saturation of 50% and 75%, the peak pull-out shear resistances varied more under

different overburden pressures. The scattered peak pull-out shear resistances in these

two series of tests might be resulted from greater disturbance during drilling in wetter

soil. This was evidenced by the difficulty experienced in drilling the holes and the

stronger vibration during drilling in the wetter soil. For tests under submerged

conditions, the submerge procedure reduced the influence of drilling disturbance so that

the results were generally constant.

The test results above suggest that there was no apparent relationship between the

pull-out shear resistance and the applied overburden pressure, which was consistent

with the observations from Cartier and Gigan (1983) and Clouterre (1991) by field

pull-out tests. The pull-out resistances under lower overburden pressure could be larger

than those under higher overburden pressure. The variation of the peak pull-out shear

resistance under different overburden pressures might have been caused by some

disturbance factors, such as misalignment of the pull-out force, variations in stress

condition of the soil and roughness of the nail surface along the nail, disturbance to the

hole surface due to drilling, etc.

5.5.2 Apparent coefficient of friction

Figure 5.13 and 5.14 show the relationship between the peak apparent coefficient of

friction μ* and applied overburden pressure for the tests in soil at degrees of saturation

- 135 -


of 38%, 50%, 75% and 98% respectively. The μ* is defined by dividing the peak

pull-out shear stress by the applied overburden pressure. The results show that the μ*

generally decreased with increase in the applied overburden pressure. The decreasing

rate of μ* decreased with the increase in overburden pressure. The μ* at higher

overburden pressure was smaller than φtan (φ is the internal friction angle of the soil)

and kept decreasing with overburden pressure, which was different from the

observations reported by Schlosser and Guilloux (1981) and Cartier and Gigan (1983).

In their observations, the μ* decreased with depth and became equal to φtan below a

certain depth. This is probably due to that different types of soil nail were used in the

tests. In the investigation performed by Cartier and Gigan (1983), driven metal nails

were used. The stresses in the soil were not substantially released during driven the

metal nails into the soil. But drill-and-grout nails were used in this project and the

stresses on the surface of the drillhole were released after drilling and did not recover

much after installing the nail. The initial normal stress acting on the nail-soil interface

was therefore much smaller in comparison with the applied overburden pressure. The

normal stress exerted on the surface of the soil nail was generally mobilized by

constrained dilatancy of the soil and was smaller than higher applied overburden

pressure. But the apparent coefficient of friction μ* was calculated by dividing the peak

pull-out shear stress by the applied overburden pressure so that it was smaller than

φtan at higher overburden pressure.

5.6 SHEAR STRESS DISTRIBUTION ON THE NAIL-SOIL

INTERFACE

The strain of the soil nail was measured by strain gauges fixed on the steel bar during

the course of pull-out. Totally four strain gauges were adhered to the steel bar at spacing

- 136 -


of about 300mm to measure the changes in strains in the steel bar and in turn the shear

stress distribution along the soil nail interface. The locations of the strain gauges are

shown in Figures 5.15 and 5.16. In the tests, the total grouted length of the steel bar was

1.2m with the length inside the box of 1.0m (0.2m of its length was within the extension

chamber). The first strain gauge was located 50mm behind the front wall of the box and

others 300mm apart.

For all the tests, the measured strains were observed to increase with the increase in

pull-out force before the peak pull-out resistance was reached. The measured strains at

the four strain gauges generally decreased in order from Strain gauge 1 to Strain gauge

4. After the peak pull-out resistance had been fully mobilized, the variation of the

strains became erratic probably due to the nonlinear response of the strain gauges at

very small strains of the large diameter steel bar subjected to residual pull-out resistance

at large displacements. Figure 5.15 shows the strain gauge responses with the pull-out

displacement for a test under overburden pressure of 40kPa and in soil at 75% degree of

saturation. Comparing with the pull-out shear stress-displacement curve shown in

Figure 5.7, it can be found that the strains measured by Strain gauges 1, 2 and 3 reached

their peak at displacements close to the displacement at which the peak shear strength

was mobilized. After the peak, they started to decrease and the strain measured by

Strain gauge 1 kept decreasing till the pull-out was finished but the strains measured by

Strain gauges 2 and 3 increased again at displacement of about 50mm. The strain

measured by Strain gauge 4 kept increasing slowly from the very start to the end of the

pull-out. The relationship between the measured strain at the peak shear strength and the

distance from the nail head is shown in Figure 5.16. The strains are observed to

decrease approximately linearly from the nail head to the nail end, which indicates a

generally uniform shear resistance along the nail length.

- 137 -


The behaviour for the curve of Strain gauge 1 is easy to understand. This strain gauge

was located close to the nail head and the pull-out force applied by the hydraulic jack

was transferred to it with little frictional loss along the nail. The strain gauge basically

measured the variation of the total pull-out force and the curve thus resembles that

pull-out shear stress-displacement curve in Figure 5.7. Figure 5.17 shows the

relationship between the pull-out force and strain measured by Strain gauge 1 and a

generally linear behaviour of the curve is observed.

The behaviour of Strain gauges 2, 3 and 4 after the peak shear stress was more complex.

Figure 5.18 illustrates the pull-out procedure. It should be noted that there was no soil

inside the extension chamber in order to maintain a constant test nail length of 1.0m

throughout the pull-out procedure. After the peak pull-out shear strength, the shear

stress on the nail-soil interface decreased, on the other hand the length of the nail inside

the soil behind Strain gauges 2, 3 and 4 increased. A theoretical analysis was conducted

to study the behaviour of these strain gauges. In the analysis, the distribution of shear

stress along the nail-soil interface was assumed to be uniform. The calculated

theoretical axial tensile forces at different pull-out displacements are shown in Figure

5.19. Compared with the corresponding strain-displacement curve in Figure 5.15, it can

be seen that the variation of the calculated axial tensile force at the location of Strain

gauge 1 and 4 is in good agreement with that of the measured strain. However, such

agreement cannot be found in the behaviour of Strain gauge 2 and 3. It is difficult to tell

the actual reason behind this phenomenon. One probable reason might be that bending

deformation was mobilized in the steel bar. If the deformation of the steel bar was

purely tensile, the strain measured by Strain gauge 3 would not exceed the strain

measured by Strain gauge 1.

- 138 -


5.7 SUMMARY

The results of the tests in soil at degrees of saturation of 38%, 50%, 75% and 98% and

under applied overburden pressures of 40kPa, 80kPa, 120kPa, 200kPa and 300kPa are

presented and discussed in this chapter. The following are the key observations from the

test results:

(a) After drilling a horizontal hole in the soil, most of the soil stresses around the hole

were released and recovery of the stresses was minimal due to grouting of the soil

nail.

(b) During pull-out, the normal stress in the soil surrounding the soil nail was increased.

The increased stress is believed to be generated by constrained dilatancy of the soil.

(c) Most of the pull-out shear stress-displacement curves exhibit a peak and post-peak

displacement softening behavior in particular for tests in soil at higher degrees of

saturation.

(d) The development of pull-out shear resistance was mainly derived from the

constrained dilatancy of the soil.

- 139 -


Earth pressure vs. time during applying overburden pressure

0

50

100

150

200

250

300

0 500 1000 1500 2000Time (min)

Earth

pre

ssur

e (k

Pa)

P-Cell 1 P-Cell 2 P-Cell 3P-Cell 4 P-Cell 5 P-Cell 6

(a)

OP=200kPaSr=38%

Vertical displacement vs. time during applying overburden pressure

0

5

10

15

20

25

30

35

40

45

50

0 500 1000 1500 2000Time (min)

Ver

tical

disp

lace

men

t (m

m) _

(b)

OP=200kPaSr=38%

Figure 5.1 – (a) Total earth pressure and (b) vertical displacement vs. time during applying overburden pressure (OP) – for overburden pressure of 200kPa and initial degree of saturation (Sr) of 38%

- 140 -


Earth pressure vs. time during drilling and grouting

0

50

100

150

200

250

300

1 10 100 1000 10000Time (min)

Earth

pre

ssur

e (k

Pa)

P-Cell 1 P-Cell 2P-Cell 3 P-Cell 4P-Cell 5 P-Cell 6

(a)

OP=200kPaSr=38%

Drill bit reached thepressure cells

Figure 5.2 – Total earth pressure vs. time during drilling and grouting – for overburden pressure of 200kPa and initial degrees of saturation (Sr) of (a) 38% and (b) 75%

Earth pressure vs. time during drilling and grouting

0

50

100

150

200

250

300

1 10 100 1000 10000

Time (min)

Earth

pre

ssur

e (k

Pa)

P-Cell 1 P-Cell 2P-Cell 3 P-Cell 4P-Cell 5 P-Cell 6

OP=200kPaSr=75%

Drill bit reached thepressure cells

P-Cell1P-Cell3P-Cell2

P-Cell4

Groutingcompleted

P-Cell6

P-Cell5

(b)

P-Cell1

P-Cell3

P-Cell2

P-Cell4

Groutingcompleted

P-Cell6

P-Cell5

- 141 -


Average earth pressure and shear stress vs. time during pull-out

020406080

100120140160180200

0 50 100 150 200 250 300 350 400Time (min)

Pres

sure

/stre

ss (k

Pa)

Earth pressureShear stress

Loading stepsPeak stress

Peak earth pressure

End of test

(a)

OP=300kPa

Sr=75%

Average earth pressure and shear stress vs. pull-out displacement

020406080

100120140160180200

0 50 100 150 200 250Pull-out displacement (mm)

Pres

sure

/stre

ss (k

Pa)

Earth pressureShear stress

Peak stress

Peak earth pressure

OP=300kPaSr=75%

End of test

(b)

Figure 5.3 – Average earth pressure and pull-out shear stress vs. (a) time and (b) pull-out displacement during pull-out – for overburden pressure of 300kPa and initial degree of saturation (Sr) of 75%

- 142 -


Changes of average earth pressure at different stages of testing

0

50

100

150

200

250

300

350

0 50 100 150 200 250 300 350Overburden pressure (kPa)

Ave

rage

ear

th p

ress

ure

(kPa

)Before drillingAfter drillingBefore pull-outAt peak resistance

Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

Figure 5.4 – Changes of average total earth pressure of P-Cells 1 to 4 at different stages of testing for tests with soil at 38% degree of saturation

Average pull-out shear stress vs. pull-out displacement

0

20

40

60

80

100

120

0 20 40 60 80 100 120 140 160 180 200 220Pull-out displacement (mm)

Ave

rage

pul

l-out

shea

r stre

ss (k

Pa)_ 40 kPa 120 kPa

200 kPa 300 kPaPeak stress

End of test


- 143 -



0

20

40

60

80

100

120

140


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

__120 kPa 200 kPa 300 kPa

Peak stress

End of test



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 40 kPa 80 kPa

120 kPa 200 kPa300 kPa

Peak stress

End of test


- 144 -



0

10

20

30

40

50

60

70

0 20 40 60 80 100 120Pull-out displacement (mm)

Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 40 kPa 80 kPa

120kPa 200 kPa300 kPa

Peak stress

End of test

Figure 5.8 – Relationship between average pull-out shear stress and pull-out displacement for tests under submerged condition (Sr≈98%)

Peak pull-out shear resistance vs. overburden pressure

0

20

40

60

80

100


Peak

pul

l-out

shea

r res

istan

ce (k

Pa)

_


- 145 -



0

20

40

60

80

100

120

0 20 40 60 80 100 120 140 160 180 200 220 240 260 280 300 320Overburden pressure (kPa)

Peak

pul

l-out

shea

r res

istan

ce (k

Pa)

__



0

20

40

60

80

100

120


Peak

pul

l-out

shea

r res

istan

ce (k

Pa) _


- 146 -



0

20

40

60

80


Peak

pul

l-out

shea

r res

istan

ce (k

Pa)

_

Figure 5.12 – Relationship between peak pull-out shear resistance and applied overburden pressure for submerged tests

0.00.20.40.60.81.01.21.41.6


Coe

ffici

ent o

f fric

tion,

μ*

Tanφ

Figure 5.13 – Relationship between the peak apparent coefficient of friction and applied overburden pressure for tests in soil at degree of saturation of 38%

- 147 -


Figure 5.14 – Peak apparent coefficient of friction vs. applied overburden pressure for tests in soil at degrees of saturation of (a) 50%, (b) 75% and (c) 98%

Tanφ'

0.00.10.20.30.40.50.60.70.8


Coe

ffici

ent o

f fric

tion,

μ*

(a)

Tanφ

0.0

0.5

1.0

1.5

2.0


Coe

ffici

ent o

f fric

tion,

μ*

(b)

Tanφ

0.00.10.20.30.40.50.60.70.80.91.0


Coe

ffici

ent o

f fric

tion,

μ*

(c)

Tanφ

- 148 -


0102030405060708090


Mic

ro-s

train

( με)

Strain 2

Strain 4

Strain 3Strain 1

50mm300mm

50mm300mm 300mm

Strain gauge 1 Strain gauge 2 Strain gauge 3 Strain gauge 4

200mmNail head

Figure 5.15 – Relationship between measured strain and pull-out displacement during pull-out for test with overburden pressure of 40 kPa and degree of saturation of 75%

0

20

40

60

80

100

0 200 400 600 800 1000 1200Distance from nail head (mm)

Mic

ro-s

train

( με)

50mm300mm

50mm300mm 300mm

Strain gauge 1 Strain gauge 2 Strain gauge 3 Strain gauge 4

200mmNail head Figure 5.16 – Strain distribution along the nail at peak pull-out stress for test with overburden pressure of 40 kPa and degree of saturation of 75%

- 149 -


Pull-out force vs. strain measured by Strain gauge 1

0

5

10

15

20

25

30

0 20 40 60 80Micro-strain (με)

Pull-

out f

orce

(kN

100

) OP=40kPa

Sr=75% Figure 5.17 – Relationship between pull-out force and strain measured by Strain gauge 1 for test with overburden pressure of 40 kPa and degree of saturation of 75%

Pullout No soil here

Strain gauges (changed locations during pull-out)

Strain gauge 1Strain gauge 4

Figure 5.18 – Illustration of pull-out procedure

- 150 -


Theoretical tensile force vs. displacement during pull-out

0

5

10

15

20

25

30


Axi

al te

nsile

forc

e (k

N) _

Strain gauge 1Strain gauge 2Strain gauge 3Strain gauge 4

OP=40kPaSr=75%

Figure 5.19 – Calculated theoretical axial tensile force vs. pull-out displacement

- 151 -

Chapter 6: Influence of degree of saturation on pull-out behaviour and resistance

CHAPTER 6: INFLUENCE OF DEGREE OF SATURATION ON

SOIL NAIL PULL-OUT BEHAVIOUR AND RESISTANCE

6.1 INTRODUCTION

In Chapter 5, the influence of overburden pressure on the pull-out resistance of the soil

nail had been discussed based on the four series of tests on soil nails in the CDG soil

under different overburden pressures and at different degrees of saturation. In this

chapter the influence of degree of saturation of soil on the pull-out behaviour and

resistance of the soil nail will be analyzed and discussed.

Among a number of factors influencing the soil nail pull-out resistance, the degree of

saturation of soil is an important influencing factor, especially for permanent soil nailed

structures. This is because the degree of saturation of the soil mass changes frequently

due to the variation of ground water table and weather conditions. The degree of

saturation of the soil can alter the suction of the soil. The pull-out resistance of a soil

nail may drop to an unsafe level in the intense rainfalls due to the increase of the degree

of saturation, that is, reduction in soil suction. However, in the previous reported

investigations, the effect of degree of saturation of soil on the nail-soil interface shear

resistance was scarcely studied.

In the literatures reviewed in Chapter 2, only a few researchers had investigated the

influence of degree of saturation of soil on the pull-out resistance of the soil nail. In the

French National Project CLOUTERRE (1991), the maximum pull-out force was found

to be reduced by more than a half when the moisture content was increased from the

- 152 -


optimum water content to the saturation moisture content. Pradhan (2003) noted from

laboratory pull-out tests with cement grouted nails in loose CDG fill at both natural

moisture content and nearly saturated conditions that only the nail-soil interface

adhesion was reduced due to high degree of saturation but the interface friction angle

kept unchanged. Chu and Yin (2005a) carried out a series of tests in CDG fill prepared

to different degrees of saturation and found that the nail-soil interface shear resistance

decreased as degree of saturation of the soil increased.

The reported investigations on the influence of degree of saturation of soil on the

pull-out resistance of a soil nail are limited. Further more, tests were only carried out in

soil from natural wet or at optimum moisture content to saturated condition but tests in

dry soil were never reported. Thirdly, in tests carried out by Chu and Yin (2005a), it was

difficult to achieve a higher degree of saturation due to leakage problem.

In this project, tests were carried out in soil at degrees of saturation from 38% to about

98%. The apparatus is waterproof and allows application of back pressure to speed up

the saturation process of the test soil. Soil samples taken after test revealed that a high

degree of saturation (about 98%) was achieved. In this chapter, the earth pressure and

pore pressure during application of back pressure, the failure patterns of soil nails at

different degrees of saturation of the soil and effect of degree of saturation of the soil on

the soil nail pull-out behaviour and resistance will be presented and discussed.

6.2 EARTH PRESSURE AND PORE PRESSURE RESPONSES

DURING SATURATING THE SOIL

Figure 6.1 shows (a) the effective earth pressure and (b) the porewater pressure

- 153 -


responses with time during de-aerating process, water infiltration and application of

back pressure for a submerged test with overburden pressure of 200kPa. A back pressure

up to 55kPa (average) was applied to accelerate the saturation of the test soil.

Measurement after completion of the pull-out test revealed that the degree of saturation

of the soil was increased to 95% from the initial value of about 84% before saturating

the soil. Suction in soil and positive porewater pressure were measured by two soil

moisture probes and one miniature porewater pressure transducer (PPT) respectively at

locations approximately 25mm to either side of the testing nail as shown in Figure 6.1 (a)

and Figure 3.21. In the longitudinal direction the two tensiometers and the PPT were

300mm and 700mm away from the nail head respectively.

In Figure 6.1 (a), the effective earth pressure is defined as the total earth pressure

measured by earth pressure cells minus the average porewater pressure measured by the

porewater pressure transducer and soil moisture probes (from negative values of about

-26kPa to a positive value of approximately 55kPa). The box was first sealed for

de-aerating with four plastic pipes connecting the four accesses on the upper side of the

side plates of the box with a vacuum pump. The vacuum pump was on for about one

hour and then the valve on the pipe connected to the vacuum pump was closed. Vacuum

was developed in the box and negative pore pressure was measured as shown in Figure

6.1 (b). Under the vacuum, water flowed into the box automatically rising upwards

inside the box from the bottom. This water inflow process lasted for about ten hours.

During this period, the pore pressure increased slowly with time. When there was no

further change in the pore pressure for about one hour, the valve connected to the

vacuum pump was reopened and the de-aerating process started again. This procedure

was repeated for several times until water was seen continuously flowing out without air

bubbles therein. The whole procedure of de-aerating and soaking the sample with water

- 154 -


lasted for more than 24 hours (about 1500 minutes).

Even with the repeated process of de-aerating and soaking, there might still be some air

inside the soil. Back water pressure was applied to dissolve the air so that a higher

degree of saturation could be achieved. Back pressure was applied in steps and the same

amount of vertical overburden pressure was applied at the same time to ensure the

effective earth pressure was constant. From Figure 6.1, it can be observed that the pore

pressure increased stepwise to an average value of about 55kPa and the effective earth

pressure only increased slightly at the end of this procedure.

6.3 VARIATIONS OF EARTH PRESSURE

The typical variations of earth pressure during the whole procedure of the pull-out test

had been presented and discussed in Chapter 5. However, the variations of earth

pressure for different series of tests in soil at different degrees of saturation were

different, which will be discussed in this chapter.

6.3.1 Decreased earth pressure immediately after grouting

As mentioned in Chapter 5, the earth pressure was observed to decrease immediately

after grouting probably due to the softening of the soil by water of the cement grout.

Figure 6.2 shows the relationship between the decrease in average earth pressure

immediately after grouting and applied overburden pressure for tests at different degrees

of saturation. The decrease in average earth pressure increased with applied overburden

pressure as had been discussed in Chapter 5. For different series of tests, the decrease in

average earth pressure was larger for tests in soil at lower degrees of saturation. For

tests in soil at lower degrees of saturation, the suction power of the soil was higher and

more water would be absorbed from the cement grout. The softening of the soil and

- 155 -


decrease in earth pressure were therefore more significant.

6.3.2 Variations of earth pressure

Figure 6.3 and Figure 6.4 show the variations of average earth pressure at different

stages of testing for all the tests with the nail grouted without pressure (gravity head

only). For all of the four series of tests, the variations of average earth pressure were

similar. The average earth pressure firstly increased to the objective value of the applied

overburden pressure during application of the overburden pressure. After drilling, the

stresses on the surface of the hole were released and the average earth pressure

substantially decreased to a residual value. A further decrease of the average earth

pressure was observed immediately after grouting probably due to softening of the soil

by water of the cement grout. After the cement grout had hardened, the average earth

pressure recovered a little bit in most of the tests due to stress redistribution. During

pull-out, the average earth pressure was observed to increase with the increase of

pull-out load. However, differences still existed for different series of tests in soil at

different degrees of saturation. For the tests in soil at 38% degree of saturation, the

average earth pressure before pull-out was smaller than that after drilling due to the

relative higher decrease in average earth pressure immediately after grouting (Figure 6.3

(a)). For other series of tests, the average earth pressure before pull-out was close to or a

little larger than that after drilling. For all the tests carried out in partially saturated soil,

the average earth pressure significantly increased at peak pull-out resistance due to

constrained dilation of the soil. However, the increased average earth pressure was

smaller for submerged tests probably due to that the dilatancy of the soil was reduced by

saturation of the soil.

- 156 -


6.4 FAILURE PATTERNS OF THE SOIL NAIL

6.4.1 Surface of the drillhole before and after pull-out

Figure 6.5 to Figure 6.8 show the surfaces of the drillholes before and after pull-out in

tests with soil at different degrees of saturation. For tests in soil at 38% and 98%

(submerged) degrees of saturation, the surfaces of the drillholes were fairly smooth

before pull-out because that less disturbance was observed during drilling. For soil at

38% degree of saturation, the hole was easy to be drilled because the soil was dry. For

submerged tests, during drilling, the soil was so wet (over 80% degree of saturation)

that could stick on the outer surface of the drill bit to prevent the drill bit from vibrating.

For soil at 50% and 75% degrees of saturation, vibration of the drill bit was observed

and small annular grooves were scratched on the surfaces of the drillholes before

pull-out. The surfaces of the drillholes after pull-out in tests with soil at 38% and 50%

degrees of saturation were generally smooth except some axial scratches along the hole.

For tests in soil at 75% degree of saturation, the surface of the drillhole after pull-out

was very rough because the failure plane was within the soil. For submerged tests,

slurry of the soil was observed on the surface of the drillhole after pull-out.

6.4.2 Failure surfaces of soil nails in the soil at different degrees of saturation

Figure 6.9 shows the nail surfaces after pull-out in tests with soil at different degrees of

saturation. The migration of shearing plane from the interface between the soil nail

surface and the surrounding soil to the interior of the soil was obvious when the soil was

getting wetter. It can be clearly seen that failure occurred mainly on the interface

between soil nail surface and the surrounding soil in tests with soil at 38% degree of

saturation. This indicates that the nail-soil interface shear strength is smaller than that of

the soil surrounding the soil nail, probably because of the strong soil suction within the

soil around the nail. For the nail in soil at 50% degree of saturation, more soil was

- 157 -


observed on the surface of the nail, which indicated that the shearing plane was

migrating into the soil. For tests in soil at 75% and 98% degrees of saturation, failure

occurred within the soil, which suggested that the nail-soil interface shear strength was

higher than that of the soil.

6.5 EFFECT OF DEGREE OF SATURATION OF THE SOIL ON


Figure 6.10 to Figure 6.14 show the relationship between average pull-out shear stress

and pull-out displacement for tests in soil at different degrees of saturation and under

overburden pressures of 40kPa, 80kPa, 120kPa, 200kPa and 300kPa respectively. It is

observed that the post-peak pull-out shear stresses generally decrease faster at higher

degree of saturation of the soil. This is probably related to the migration of shearing

plane from the interface between the soil nail surface and the surrounding soil to the

interior of the soil when the soil is getting wetter. For the failures within the soil, the

pull-out stress-displacement behaviour showed obvious displacement softening as

generally shown in the behaviour of dense dilative soil. For the failures on the nail-soil

interface, there was small and slow reductions of the post-peak pull-out shear resistance,

i.e., the displacement softening was not the dominating behaviour of the dilative soil

subjected to shearing against a hard surface.

From Figure 6.10 to Figure 6.14, it can be observed that, for all the applied overburden

pressures, the peak pull-out shear strengths in tests with soil prepared to 50% and 75%

degrees of saturation were greater than those for saturated tests and tests with soil

prepared to 38% degree of saturation. For soil at degree of saturation of 75%, the

moisture content is close to the optimum moisture content. The peak pull-out shear

strength in tests at degree of saturation of 75% was about 2 times that in the saturated

- 158 -


tests. The pull-out displacement at peak pull-out shear strength increased with decrease

in degree of saturation of the soil. The displacement at peak pull-out shear strength in

tests at degree of saturation of 75% was about 2 times as much as that in the saturated

tests. These results are similar to the results presented in CLOUTERRE (1991) which

showed that the maximum pull-out force was increased by 2 times when the moisture

content was decreased from saturation to the optimum water content and the

displacement corresponding to this maximum force was increased by 3 times. The

decrease in pull-out resistance with degree of saturation from the optimum moisture

content to the saturated condition was also observed by Chu and Yin (2005a) and

Pradhan (2003). Pradhan (2003) considered that this was because of the decrease in soil

cohesion.

Figure 6.15 to 6.17 show the relationship between the peak pull-out shear resistance and

degree of saturation with soil under different vertical pressures. From the plotted trend

lines in the figures, it can be observed that the peak pull-out shear strength firstly

increased and then decreased with increase in degree of saturation of the soil. In the

tests under the overburden pressure of 80kPa, only the decrease in pull-out resistance

when the degree of saturation of the soil increased from 75% to 98% was observed

because that the tests in soil at 38% and 50% degrees of saturation had failed. It is

believed that if the soil is dry, water is more readily to be absorbed by the soil due to a

higher suction power. This may cause more contraction of the cement grout thus

reducing the bond strength between the nail surface and the surrounding soil. In the tests

with soil at 38% degree of saturation, as mentioned in the above paragraphs, further

decrease of earth pressure was observed immediately after grouting and the earth

pressure did not recover after the cement grout had hardened. This is a possible

indication of the contraction of the cement grout. The actual reason is unknown and

- 159 -


needs to be found out by further tests and analyses. In this respect, tests conducted using

pressure grouting method to improve the contact between the nail surface and the

surrounding soil may throw some light into it. As in the submerged tests, the decrease in

soil cohesion could be one of the reasons for the decrease in pull-out resistance. The

decrease in dilatancy of the soil could be another reason. In Figures 6.3 and 6.4, it is

observed that the increases in earth pressures at peak pull-out resistance for submerged

tests were much smaller than those in tests with soil at other degrees of saturation.

The peak pull-out shear strength for tests at degree of saturation of 75% was about 2

times that for saturated tests. Some of the tests in soil at 50% degree of saturation were

even higher (Figures 6.15 to 6.17). This indicates that the effect of degree of saturation

on soil nail pull-out resistance is significant and should be carefully addressed in design

of the soil nailing system.

6.6 SUMMARY AND MAJOR FINDINGS

The effect of degree of saturation of the soil on the soil nail pull-out behaviour and

resistance is discussed in this chapter. The following are the key observations from the

test results:

(a) The migration of shearing plane from the interface between the nail surface and the

surrounding soil to the interior of the soil was observed with the increase in degree

of saturation of the soil.

(b) The displacement at peak pull-out shear resistance for saturated tests was smaller

than those for partially saturated tests. The decreasing rates of the post-peak pull-out

shear stress increased with the increase in degree of saturation of the soil.

(c) The peak pull-out shear resistance varies with different degrees of saturation of the

soil, with higher resistances at the degrees of saturation of 50% and 75%.

- 160 -

cesulijun

Note

MigrationConfirmed set by cesulijun

cesulijun

Note

MigrationConfirmed set by cesulijun


(d) The increases in earth pressures at peak pull-out resistance for submerged tests were

much smaller than those in tests at other degrees of saturation. This indicates that

the soil dilatancy was reduced when the soil was saturated and as a result the

pull-out resistance decreased.

It should be emphasized that the number of tests carried out was limited and only four

different degrees of saturation were tested. It is premature to draw any conclusion on the

precise correlation between the degree of saturation of the soil and the pull-out

resistance of soil nail. Further research in this area is needed.

- 161 -


Effective earth pressure vs. time during applying back pressure

-50

0

50

100

150

200

250

0 500 1000 1500 2000 2500 3000 3500 4000Time (min)

Effe

ctiv

e ea

rth p

ress

ure

(kPa

) _


2 1

346 5

PPT T1,T2

(a)

OP=200kPaSubmerged

Pore pressure vs. time during applying back pressure

-60

-40

-20

0

20

40

60

80

0 500 1000 1500 2000 2500 3000 3500 4000Time (min)

Pore

pre

ssur

e (k

Pa)

PPT Tensiometer 1 Tensiometer 2

“-” for suction and “+” forpositive pore water pressure

Applying back water pressure

Air release andWater infiltration

(b) Figure 6.1 – (a) Effective earth pressure and (b) porewater pressure vs. time during saturating the soil – for a submerged test with overburden pressure of 200kPa

- 162 -


Decreased average earth pressure immediately after grouting

0

5

10

15

20

25

30


Dec

reas

ed a

vera

ge e

arth

pre

ssur

e (k

Pa) _ 38% 50% 75%

Figure 6.2 – Relationship between the decrease in average earth pressure immediately after grouting and applied overburden pressure for tests at different degrees of saturation

- 163 -



0

50

100

150

200

250

300

350


Ave

rage

ear

th p

ress

ure

(kPa

)Before drillingAfter drillingBefore pull-outAt peak resistance

Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

(a)


0

50

100

150

200

250

300

350


Figure 6.3 – Changes of average earth pressure at different stages of testing for tests with soil at (a) 38% and (b) 50% degrees of saturation

Ave

rage

ear

th p

ress

ure

(kPa

) _

Before drillingAfter drillingBefore pull-outAt peak resistance

Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

(b)

- 164 -



0

50

100

150

200

250

300

350


Figure 6.4 – Changes of average earth pressure at different stages of testing for tests with soil at (a) 75% and (b) 98% degrees of saturation

Ave

rage

ear

th p

ress

ure

(kPa

) _Before drillingAfter drillingBefore pull-outAt peak resistance

Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

(a)


0

50

100

150

200

250

300

350


Ave

rage

ear

th p

ress

ure

(kPa

) _


Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

(b)

- 165 -


(a) (b) Figure 6.5 – Surfaces of the drilllhole before and after pull-out in a test with soil at 38% degree of saturation

- 166 -


(a)

(b) Figure 6.6 – Surfaces of the drillhole before and after pull-out in a test with soil at 50% degree of saturation

- 167 -


(a)

(b) Figure 6.7 – Surfaces of the drillhole before and after pull-out in a test with soil at 75% degree of saturation

- 168 -


(a) (b) Figure 6.8 – Surfaces of the drillhole before and after pull-out in a test with soil at 98% degree of saturation

- 169 -


(a)

(b)

(c) (d) Figure 6.9 – Nail surfaces after pull-out in tests with soil at degrees of saturation of (a) 38% (b) 50% (c) 75% and (d) 98%

- 170 -



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

38 % 75 %98 %Peak Stress

Overburden Pressure=40kPa

Figure 6.10 – Relationship between average pull-out shear stress and pull-out displacement for tests at different degrees of saturation under overburden pressure of 40kPa


0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

__

75 % 98 %Peak Stress



- 171 -



0

20

40

60

80

100

120

0 20 40 60 80 100 120 140 160 180 200 220 240Pull-out displacement (mm)

Ave

rage

pul

l-out

shea

r stre

ss (k

Pa)__ 38 % 50%



Figure 6.12 – Relationship between average pull-out shear stress and pull-out displacement for tests at different degrees of saturation and overburden pressure of 120kPa


0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 38 % 75 %

98 % 50%Peak Stress


Figure 6.13 – Relationship between average pull-out shear stress and pull-out displacement for tests at different degrees of saturation and overburden pressure of 200kPa

- 172 -



0

20

40

60

80

100

120

140


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 38 % 50%




Peak pull-out shear resistance and degree of saturation

0

20

40

60

80

100

0 20 40 60 80 100 120Degree of saturation (%)

Peak

resis

tanc

e (k

Pa)

__

Figure 6.15 – Relationship between peak pull-out shear resistance and degree of saturation with overburden pressure of 40kPa

- 173 -



0

20

40

60

80


Peak

resis

tanc

e (k

Pa)

__

20

(a)


0

20

40

60

80

100


Peak

resis

tanc

e (k

Pa)

__

(b) Figure 6.16 – Relationship between peak pull-out shear resistance and degree of saturation with overburden pressures of (a) 80kPa and (b) 120kPa

- 174 -



0

20

40

60

80

100

0 20 40 60 80 100Degree of saturation (%)

Peak

resis

tanc

e (k

Pa)

__

120

(a) Peak pull-out shear resistance and degree of saturation

0

20

40

60

80

100

120


Peak

resis

tanc

e (k

Pa)

(b) Figure 6.17 – Relationship between peak pull-out shear resistance and degree of saturation with overburden pressure of (a) 200kPa and (b) 300kPa

- 175 -

Chapter 7: Effect of grouting pressure on pull-out behaviour and resistance

CHAPTER 7: EFFECT OF GROUTING PRESSURE ON SOIL

NAIL PULL-OUT BEHAVIOUR AND RESISTANCE

7.1 INTRODUCTION

In common practice of the soil nail construction, gravity or low pressure grouting is

normally adopted. Some initial results from field pull-out tests on soil nails grouted

under high pressure have indicated that the grouting pressure contributed a lot to the

pull-out resistance. The pressure grouting is a cost effective method for increasing the

soil nail pull-out resistance and in turn improving the performance of the nailed

structure. Studies on the significance of the pressure grouting are common for anchors,

piles and ground improvement, but limited for soil nails. Haberfield (2000) presented a

study on the prediction of the initial normal stress in piles and anchors constructed using

expansive cements. Moosavi et al. (2005) reported a study on the bond of cement

grouted bars under pressure. Soga et al. (2005) presented results of a laboratory

investigation of multiple grout injection into clay. Yeung et al. (2005) carried out field

pull-out tests on Glass Fiber Reinforced Polymer (GFRP) pipes with cement pressure

grouting in a CDG soil slope in Hong Kong and observed a significant increase in the

pull-out resistance due to the pressure grouting. Study on the influence of grouting

pressure on soil nail pull-out resistance together with the other factors, such as the

overburden pressures and soil conditions, etc., is limited, especially for soils in Hong

Kong.

In this project, a total of four pull-out tests have been carried out on soil nails grouted

with grouting pressures of 80kPa and 130kPa respectively. The degree of saturation of

- 176 -


the soil was 50% and applied overburden pressures were 80kPa and 200kPa respectively.

In this chapter, the results of these tests will be presented and the influence of grouting

pressure on the soil nail pull-out behaviour and resistance will be discussed.

7.2 VARIATIONS OF EARTH PRESSURES

7.2.1 Variations of earth pressures during drilling and pressure grouting

As discussed in Chapter 5, in current practice of soil nail installation, the hole-drilling

may cause stress release in soil around the drillhole and this would affect the earth

pressure acting on the soil nails. Reasonable simulation of the stress release and

examination of the earth pressure changes are therefore important. Figure 7.1 (a) shows

the measured earth pressures at the six earth pressure cells versus time during drilling

and pressure grouting for a test in soil at 50% degree of saturation under overburden

pressure of 200kPa and with grouting pressure of 130kPa. Because the grouting time

was short comparing with the time for the whole procedure, the variations of earth

pressures during pressure grouting were not clear in Figure 7.1 (a). Therefore the

variations of earth pressures measured by earth pressure cells 1 to 4 during pressure

grouting are shown in Figure 7.1 (b). The variations of earth pressures during drilling

were similar to those in tests with nails grouted without pressure (gravity head only) as

discussed in Chapter 5 and it is not necessary to repeat the discussion here.

After the drillhole was formed, a high yield steel bar with 40mm in diameter was

installed in the centre of the hole and pressure grouting was conducted. Figure 7.1 (b)

shows rapid increases in the earth pressures at the four earth pressure cells due to the

action of the pressure grouting. The earth pressures increased quickly from initial values

of about 50kPa to 70kPa to peak values of about 140kPa to 160kPa in less than ten

minutes. Immediately after the peak values were achieved, the earth pressures started to

- 177 -


decrease and finally reached values approximating to their original values before

pressure grouting. The reasons for this are explained as follows.

(a) First, the cement grout inside the plastic grouting tube set with time quickly,

leading to a significant increase in friction between the cement grout and the tube.

As a result, the applied grouting pressure in the pressure grouting cylinder could not

be totally transmitted to the cement grout inside the drillhole due to friction loss.

(b) Second, as hardening developed with time, the cement grout in the plastic grouting

tube could not move much or would not move at all.

(c) On the other hand, the viscosity of the grout inside the drillhole increased with time,

preventing further grout from entering the drillhole through the plastic tube.

(d) In fact, the factors (a), (b) and (c) above caused the grouting pressure inside the

drillhole to decrease with time elapsing.

(e) At the same time, the water inside the cement grout was absorbed by the soil and as

such the water pressure in the cement grout was dissipated into the soil. This was

similar to the dissipation of excess pore water pressure in a saturated soil.

(f) It is considered that the factor in (e) together with the decrease in the grouting

pressure in (d) caused the cement grouting pressure inside the drillhole to drop

rapidly shortly after grouting. After the cement grout had solidified in the last

period of the soil nail installation, the earth pressures at earth pressure cells 1 to 4

increased slightly, probably due to re-arrangement of the soil particles.

It will be seen that, even with this short duration of the pressure increase, the soil nail

pull-out resistance has been significantly increased. The reasons for this are given in the

remaining sections of this chapter.

- 178 -


7.2.2 Variations of earth pressures during the whole period of testing

Figure 7.2 shows the changes of average earth pressures measured by earth pressure

cells 1 to 4 at different stages of testing for the tests in soil at 50% degree of saturation

with grouting pressure of (a) 80kPa and (b) 130kPa respectively. The variations of the

average earth pressures are similar to those in tests with nail grouted by gravity head

only as discussed in Chapter 5 and 6. Compared with Figure 6.3 (b), it can be seen that

the average earth pressures before pull-out (after the cement grout had hardened) for

tests with nail grouted at pressure of 80kPa, 130kPa and 0kPa (gravity head only) were

all close to each other. This is because that the pressure grouting did not permanently

increase the earth pressure as has been discussed in the above paragraphs. However, at

peak pull-out resistance, the average earth pressures in pressure grouting tests were

much higher than those in gravity grouting tests. For most of the pressure grouting tests

(3 out of 4), the average earth pressures at peak pull-out resistance were even higher

than the applied overburden pressure.

7.3 FAILURE PATTERNS OF THE SOIL NAIL

Figure 7.3 shows the soil nail surfaces after pull-out in tests with different grouting

pressures. It can be observed that the soil adhered to the soil nail surface increased with

the increase in grouting pressure. For the nail grouted without pressure, there were some

soil loosely adhered to the nail surface after pull-out and no scratch of the soil was

observed. This indicates that the contact between the surface of the soil nail and the

surrounding soil was not very tight, probably because of the loss of water of the cement

grout due to soil suction within the soil around the nail. For the nail grouted at grouting

pressure of 80kPa, more soil was observed to tightly adhere to the nail surface after

pull-out and some slight scratches parallel to the axis of the soil nail were observed.

This indicates that the grouting pressure improved the contact between the nail surface

- 179 -


and the surrounding soil. For the nail grouted at grouting pressure of 130kPa, the

amount of the soil adhered to the nail surface was observed to have further increased

compared with the nail grouted at grouting pressure of 80kPa. The soil was observed to

tightly adhere to the nail surface and slight scratches parallel to the axis of the soil nail

were also observed. This indicates that the increase in grouting pressure further

improved the contact between the nail surface and the surrounding soil. From the failure

patterns of the soil nail, it can be concluded that the increase in grouting pressure

improved the contact between the nail surface and the surrounding soil even though the

grouting pressure just temporarily increased the earth pressure in soil surrounding the

soil nail.

7.4 INFLUENCE OF GROUTING PRESSURE

The results from the four soil-nail pull-out tests with grouting pressures of 80kPa and

130kPa under overburden pressures of 80kPa and 200kPa together with a gravity

grouting test under overburden pressure of 200kPa are interpreted together here. These

results are used to examine the influences of grouting pressures on the soil nail pull-out

behaviour and resistance. The gravity grouting test under the overburden pressure of

80kPa had failed and is not included here.

Figure 7.4 (a) shows the measured average pull-out shear stress versus pull-out

displacement for tests with grouting pressures (GP) of 80kPa and 130kPa under the

same overburden pressure (OP) of 80kPa and in soil at degree of saturation (Sr) of 50%.

It is clear that the average pull-out shear stress increased with the applied grouting

pressure. Figure 7.4 (b) shows the results of those tests in soil at the same Sr with GP of

0kpa (gravity grouting), 80kPa and 130kPa under a different OP of 200kPa. It can be

noted that the average pull-out shear stress increased with the grouting pressure.

- 180 -


Furthermore, for the test with a grouting pressure of 130kPa, the curve shows clearly a

peak and post-peak displacement-softening behaviour. In contrast, for the test with a

grouting pressure of 80kPa, the curve indicates a relatively ductile behaviour without a

distinct peak in shear stress. The test with zero grouting pressure also exhibits a ductile

behaviour but with much lower pull-out shear stress. The differences in the

load-displacement responses of different tests reflect the differences of the grouting

pressure influences on the surrounding soil and the soil-nail interface among the three

tests. Under higher grouting pressure, the amount of soil adhered to the nail surface

increased which indicates that the failure surface was migrating from the nail-soil

surface to the interior of the soil as the grouting pressure increased. For nail grouted

with grouting pressure of 130kPa, the failure could be considered to occur within the

soil. Further more, the soil surrounding the nail was further compressed under the high

grouting pressure of 130kPa and became denser. The stress-displacement behaviour of

dense sandy soil is generally displacement-softening during shearing.

From Figure 7.4, the influences of the grouting pressures on the pull-out resistance of

soil nails are clearly observed. The increase in the soil nail peak pull-out resistance due

to the increase in grouting pressure may be explained as follows:

(a) First, after drilling, the soil at and close to the internal surface of the hole had been

disturbed and thus was loosened, and in the meantime soil arching was developed

to support the soil above the hole. Under the effect of grouting pressure, the cement

grout inside the hole compacted and densified the soil within the disturbed zone

around the hole. Depending on the magnitude of grouting pressure, the compacted

soil could be stronger than that before grouting.

(b) Second, some of the cement grout might have infiltrated into the soil around the

drillhole and increased the bond strength of the nail-soil interface and the cohesion

- 181 -


of the soil. The higher the grouting pressure, the more infiltration of the cement

grout into the soil would be.

(c) Third, the drillhole formation could have been rough because of the unevenness of

the drill bit teeth, resulting in a rough surface of the grouted nail. The higher the

grouting pressure, the greater the roughness of the surface of the grouted nail would

be. This resulted in a greater contact area with soil and soil dilation during shear

movement between the nail and the surrounding soil.

(d) Though the grouting pressure had dropped after grouting, the processes in (a), (b)

and (c) were irreversible. One or a combination of the factors explained in (a), (b)

and (c) then caused the pull-out resistance to increase.

Figure 7.5 shows plots of the average pull-out shear stress versus pull-out displacement

of the test nails in soil at Sr=50% under OP= 80kPa and 200kPa with grouting pressures

of (a) 80kPa and (b) 130kPa respectively. It can be observed that, for pressure grouting

tests, the average pull-out shear stress-displacement behaviour and peak pull-out shear

resistance were related to the overburden pressure to a certain extent. However, the

residual pull-out resistance at large pull-out displacement was not directly related to the

overburden pressure. How the overburden pressure affects the curves and the pull-out

resistance with the presence of grouting pressure needs more test results and further

study.

Figure 7.6 shows plots of (a) peak shear strength, (b) shear stress at displacement of

100mm, and (c) shear stress at displacement of 200mm versus grouting pressure under

the overburden pressure OP=80kPa and 200kPa. For the test nail prepared under a

grouting pressure of 80kPa, the peak pull-out shear strengths for the two overburden

pressures were nearly the same. However, for the one prepared under a higher grouting

- 182 -


pressure of 130kPa, the pull-out shear strengths for OP=200kPa was substantially

higher than that for OP=80kPa. This illustrates that the peak pull-out shear strength of

the nail was less dependent on the overburden pressure when the grouting pressure was

low, but the peak shear strength increased with increasing overburden pressure when the

grouting pressure was high. The probable explanation is that when the grouting pressure

was low, the disturbed soil around the drill hole was still loose and the soil arching

effect existed to a certain extent. When the grouting pressure was high, the loosened soil

was well compacted by the grout; and as a result, the strength of the soil increased back

to or even higher than that before the drilling. This led to a change in the stress

condition in the soil around the hole and the soil arching effect no longer existed. As

such, most of the vertical stress above the nail was transferred to the nail surface. At the

displacement of 100mm, the pull-out stress for the two tests under overburden pressure

of 80kPa and 200kPa with grouting pressure of 80kPa were equal to each other and the

difference between the pull-out stresses for the two tests with grouting pressure of

130kPa was reduced. At the displacement of 200mm, the pull-out resistances for tests

under 80kPa and 200kPa were close to each other with both of the two grouting

pressures. This indicates that the influence of grouting pressure was reduced at large

pull-out displacement.

7.5 SUMMARY AND CONCLUSIONS

The results for the pressure grouting tests are presented and the influence of the grouting

pressure on the pull-out behaviour and resistance of the soil nail is discussed in this

chapter. Based on the discussion, the following observations and conclusions may be

presented here:

(a) The pressure grouting device, grouting pressure measurement and procedures are

simple, practical, and reliable for studying pressure grouted soil nails using a

- 183 -


laboratory pull-out box.

(b) The grouting pressure inside the hole increased quickly, but then dropped rapidly due

to hardening of the cement grout and shrinkage of grout volume because of water

seepage from the cement grout to the surrounding soil (accompanied by dissipation

of the excess water pressure inside the grout and absorbing of water in the grout by

the dryer surrounding soil).

(c) At peak pull-out resistance, the earth pressures measured by earth pressure cells

close to the nail for pressure grouting tests increased to much higher values than

those for gravity grouting tests. For most of the tests, the earth pressure at peak

pull-out resistance was even higher than the applied overburden pressure. This is

probably because that the contact between the soil nail surface and the surrounding

soil was improved by the grouting pressure so that the effect of the constrained

dilation of soil was strengthened.

(d) The amount of soil adhered to the nail surface after pull-out increased with the

increase in grouting pressure.

(e) The average peak pull-out shear resistance of the soil nail increased almost linearly

with the increase in grouting pressure in the present study.

It shall be noted that the amount of the pressure grouting tests was limited and only one

type of CDG soil at the degree of saturation of 50% was tested in this study. More soil

nail pull-out tests on more types of soil at different degrees of saturation shall be carried

out to examine the influences of the grouting pressure. The present study may serve as a

good basis for further investigations in this area.

- 184 -


Earth pressure vs. time during drilling and pressure grouting

0

50

100

150

200

250

300

350

1 10 100 1000 10000Time (min)

Earth

pre

ssur

e (k

Pa)


Drilling

Before grouting After groutingDuring grouting

2 1

346 5

(a)

Earth pressure vs. time during pressure grouting

0

20

40

60

80

100

120

140

160

180

920 925 930 935 940 945 950 955 960Time (min)

Earth

pre

ssur

e (k

Pa)

(b) Figure 7.1 – Earth pressure vs. time during (a) drilling and grouting and (b) pressure grouting only – for a test in soil at 50% degree of saturation under overburden pressure of 200kPa and with grouting pressure of 130kPa

- 185 -



0

50

100

150

200

250

300

0 50 100 150 200 250Overburden pressure (kPa)

Ave

rage

ear

th p

ress

ure

(kPa

) _Before drillingAfter drillingBefore pull-outAt peak resistance

Before drilling

At peak resistance

After drilling

Before pull-out

2 1

34

(a)


0

50

100

150

200

250

300

0 50 100 150 200 250Overburden pressure (kPa)

Ave

rage

ear

th p

ress

ure

(kPa

) _


Before drillingAt peak resistance

After drilling Before pull-out

2 1

34

(b) Figure 7.2 – Variation of average total earth pressure for tests in soil at 50% degree of saturation with grouting pressure of (a) 80kPa and (b) 130 kPa

- 186 -


(a)

(b)

(b) (c) Figure 7.3 – Failure surfaces of the soil nails for tests in soil at 50% degree of saturation with grouting pressures of (a) 0kPa (gravity head only) (b) 80kPa and (c) 130kPa

- 187 -



0

20

40

60

80

100

120

140

160

180


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

__80 kPa 130 kPa

Peak Stress

(a) Average pull-out shear stress vs. pull-out displacement

020406080

100120140160180200220


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 0 kPa 80 kPa 130 kPaPeak Stress

(b) Figure 7.4 – Average pull-out shear stress vs. pull-out displacement for tests in soil at 50% degree of saturation with grouting pressures (GP) of 0, 80, 130kPa and under overburden pressures of (a) 80kPa and (b) 200kPa

- 188 -



020406080

100120140160180200220


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 80 kPa 200 kPaPeak Stress

(a)


020406080

100120140160180200220240


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _ 80 kPa 200 kPa

Peak Stress

(b) Figure 7.5 – Average pull-out shear stress vs. pull-out displacement for tests in soil at 50% degree of saturation under overburden pressures of 80kPa and 200kPa with grouting pressures of (a) 80kPa and (b) 130kPa

- 189 -


Peak pull-out shear resistance vs. grouting pressure

020406080

100120140160180

0 20 40 60 80 100 120 140Grouting pressure (kPa)

Peak

resis

tanc

e (k

Pa) OP=80kPa OP=200kPa

(a)

Pull-out shear stress at 100mm vs. grouting pressure

020406080

100120140160180


Pull-

out s

hear

stre

ss (k

Pa)

OP=80kPa OP=200kPa

(b)

Pull-out shear stress at 200mm vs. grouting pressure

020406080

100120140160180


Pull-

out s

hear

stre

ss (k

Pa)

OP=80kPa OP=200kPa

(c) Figure 7.6 – (a) Peak shear resistance, (b) shear stress at displacement of 100mm and (c) shear stress at displacement of 200mm versus grouting pressure for tests in soil at 50% degree of saturation under overburden pressure OP=80kPa and 200kPa

- 190 -


CHAPTER 8: NUMERICAL SIMULATION OF PULL-OUT

TESTS

8.1 INTRODUCTION

Physical modeling can obtain direct measured, accurate and reliable results on the

influences of certain parameters under strictly controlled test conditions. However, a

parametric study using a physical model is expensive, time consuming, and, sometimes,

impossible such as requirements of space and special instrumentation. Numerical

modeling can overcome these difficulties. In numerical simulation, a parametric study

can be easily conducted by simply changing one parameter at one time with other

parameters unchanged. Numerical modeling methods, especially the finite element

method, have been successfully used in simulation of the soil nailing system as in the

literature reviewed in Chapter 2. The finite element method is a numerical procedure,

which can be used to obtain solutions to a large class of engineering systems, including

analysis of stress, heat transfer, fluid flow and electromagnetism problems. In

geotechnical engineering, conventional methods for design and stability analysis are the

limit equilibrium analysis methods. These methods are generally based on a series of

assumptions and can only examine the overall resistance and stability of geotechnical

structures. Limit equilibrium analysis methods cannot predict the deformation and stress

(strain) distribution in the structures. The finite element method has the advantages of

simulating the deformation, stress (strain) distribution, and failure of geotechnical

structures.

In this project, ABAQUS, a powerful finite element (FE) program, is used to simulate

- 191 -

Chapter 8: Numerical simulation of the pull-out tests

the pull-out tests carried out in this project and to conduct further parametric studies. A

three dimensional (3-D) finite element model is established for pull-out tests and is

verified by comparing simulated results with measured data. If the agreements between

the simulated results and the test results are good, this indicates that this 3-D FE model

is applicable to other simulations. Using this model, the influences of the grouting

pressure and the dilation angle of the shearing zone on the soil nail pull-out resistance

are studied. The method for simulating the shearing plane, description of the model and

simulated results are presented and discussed in this chapter.

8.2 SIMULATION OF THE SHEARING PLANE

There are normally three types of methods for simulating the interface problem in

ABAQUS (ABAQUS6.3-1 Online Doc., 2002). The first method is to simulate an

interface using zero thickness interface elements, such as the Gap elements and

Tube-to-tube contact elements. The Gap elements can be used to simulate the contact

and separation between two nodes and a frictional coefficient can be defined to simulate

the tangential interaction behaviour. The Tube-to-tube contact elements can be used to

model the finite-sliding interaction along a slide line between two pipelines or tubes

where one tube lies inside the other or between two tubes or rods that lie next to each

other. This type of elements can be used only with first-order pipe, beam, or truss

elements and does not consider deformations of the tube or pipe cross-section. The

second method is using surface based contact to simulate the interaction between two

contact surfaces. Both normal and tangential properties can be defined for the contact

pair. The tangential properties of the contact pair can be smooth (no friction), rough (no

slide), frictional, etc. However, cohesion cannot be defined for the tangential properties

of a contact pair. The third method is to use a thin layer of continuum elements to

simulate the interface. The normal and tangential behaviour of the interface are both

- 192 -


simulated by the material properties of the continuum elements of the thin layer.

The Gap elements are similar to the surface based contact method which cannot

simulate the cohesion effect between the two surfaces. Further more, convergence

problems are often encountered when using these two methods. Tube-to-tube contact

elements are restricted to some particular problems and cannot simulate general contact

between two surfaces. As a result, the method of using a thin layer of continuum

elements is used to simulate the shearing plane in the soil nail pull-out test. A linear

elastic-perfectly plastic model using the Mohr-Coulomb failure criterion is adopted to

define the material properties of continuum elements in the shearing zone.

8.3 DESCRIPTION OF THE FINITE ELEMENT MODEL

8.3.1 Mesh and boundary conditions

A 3-D FE model is developed to simulate the pull-out tests. The dimensions of the

model are 1.0m in length, 0.3m in width and 0.81m in height. The dimensions of the test

soil sample in the pull-out tests are 1.0m in length, 0.6m in width and 0.8m in height.

The 10mm thickness on the top of the FE model is used to simulate the wooden board

placed on the test soil during the tests. Only a half of the physical model is simulated

because of the symmetry of the geometry, load and boundary conditions of the tests.

The nail is 0.1m in diameter and 1.2m in length and is formed with its center at the level

of 0.315m in height. A thin layer of material with the thickness of 4.0mm surrounding

the nail is used to simulate the shearing zone. The mesh of the model is shown in Figure

8.1. The mesh is fine around the nail and becomes coarser as the distance from the nail

surface increases. The soil, the nail and the shearing zone are all made up of C3D8R

elements, a type of 8-node linear brick elements with reduced integration. There is no

interface between the nail and the surrounding soil. As the soil nail is being pulled out,

- 193 -


the relative sliding of the nail and the surrounding soil is simulated by the plastic flow

of the shearing zone. During calculation unsymmetric matrix is used because of the high

nonlinearity of the problem.

The displacements on both left and right surfaces of the model in direction-1 (direction

perpendicular to these surfaces) are fixed at zero to simulate the symmetrical plane and

the side boundary of the box. The displacements at the bottom of the model in

direction-3 (vertical direction) are fixed at zero to simulate the bottom support of the

box. During simulating the application of the overburden pressure and formation of the

drillhole, the displacements on the front and back surfaces of the model in direction-2

are fixed at zero to simulate the front and back boundaries of the box. During simulation

of the pull-out process, the direction-2 displacements constraints at locations of the two

drilling accesses in the front and back plates of the pull-out box are released. A nonzero

displacement boundary condition is applied on the nail head to simulate the pull-out

displacement. The load and boundary conditions of the model before and during

pull-out are shown in Figures 8.2 and 8.3 respectively.

8.3.2 Procedure of the simulation

In order to simulate the actual procedure of the pull-out tests, the formation of the

drillhole and installation of the nail need to be simulated. The excavation of the drillhole

can be accomplished by simply removing the elements at the location of the drillhole.

However, upon installation of the nail, we will face two difficulties. One is that the

material properties are global model variables which cannot be changed during

calculation. Another is that deformations have been generated in elements surrounding

the drillhole due to removal of elements so that activation of the nail elements will

result in superposition of elements. These can be overcome by simulating the drilling

and grouting separately but step-by-step in two models. The stresses obtained from the

- 194 -


first model are input into the second model as initial stresses and the deformations

induced by the drilling procedure are then eliminated by equilibrium between the initial

stresses and the applied loads. The elements forming the soil nail can be activated and

assigned with nail properties in the second model to simulate the installation of the nail.

Therefore the application of the overburden pressure, formation of the drillhole and

pull-out of the soil nail were simulated in three separate models. The detailed procedure

of the simulation is as follows:

(a) In the first model, all the elements are assigned with the soil properties to simulate

the test soil compacted in the pull-out box. The elements forming the extension part

of the nail behind the box is removed because only the soil in the pull-out box is

simulated. On all the vertical surfaces and the bottom surface, the corresponding

displacements are constrained as has discussed in the forgoing paragraphs. The

gravity load is applied to the entire model and pressure load is applied on the top

surface of the model. After the calculation of this model has completed, the stresses

in all the elements are exported.

(b) In the second model, the same loading and boundary conditions as in the first model

are used and the calculation is accomplished in two steps. In the first step, the

stresses exported from the first model are imported as initial stresses of the elements.

At the end of this step, equilibrium between the initial stresses and the applied loads

is established. In the second step, the elements forming the nail are removed to

simulate the drilling process. At the end of this step, the stresses in all the elements

are exported again to be used as initial stresses for the third model.

- 195 -


(c) In the third model, the elements forming the nail and the shearing zone are assigned

with their own properties respectively. This model includes four steps. In the first

step, the elements forming the nail are removed and equilibrium between the applied

loads and the initial stresses imported from the second model is established. In the

second step, the elements forming the nail are activated without stress. In the third

step, gravity load is applied to the nail elements to simulate the grouting procedure.

In the final step, displacements constraints at locations of the two drilling accesses

are released and a pull-out displacement is applied on the nail head to simulate the

pull-out of the nail.

8.3.3 Material properties

The constitutive model used for the nail is an isotropic linear elastic model. The

material properties of the nail are calculated considering a combination of the known

stiffness of the steel bar and the cement grout. The density, Young’s Modulus and

Poisson’s ratio for the nail are calculated as 2.0Mg/m3, 16.37GPa and 0.28 respectively.

For the soil and the material in the shearing zone, an isotropic linear elastic-perfectly

plastic model using the Mohr-Coulomb failure criterion is adopted. Non-associate flow

rule is used. The model is expressed in terms of stress invariants in ABAQUS

(ABAQUS/Standard online Doc., 2002) as follows:

Mean total stress,

3321 σσσ ++

=p (8.1)

Mises equivalent stress,

)(23 S:S=q (8.2)

The third invariant of deviatoric stress

- 196 -


31

)( 29 S:SS ⋅=r (8.3)

where S is the deviatoric stress tensor and defined as IσS p−= ; is stress tensor; pI

is hydrostatic stress tensor; and I is unit tensor.

σ

Therefore, the Mohr-Coulomb model is defined by:

0tanRF =−+= cpqmc φ (8.4)

where

φπθπθφ

φθ tan3

cos31

3sin

cos31),(R ⎟

⎠⎞

⎜⎝⎛ ++⎟

⎠⎞

⎜⎝⎛ +=mc (8.5)

φ is the slope of the Mohr-Coulomb yield surface in the p-Rmcq stress plane, which is

commonly referred to as the friction angle of the material; c is the cohesion of the

material; and θ is the deviatoric polar angle defined as . 3)/()3cos( qr=θ

The flow potential G is chosen as:

( ) ( ) ψψ tanRtanG 220 pqac mw ++= (8.6)

where

( φθθ

θθ π ,R45cos)1(4)12(cos)1(2

)12(cos)1(4),(R 32222

222

mcmweeeee

eee−+−−+−

−+−= ) (8.7)

ψ is the dilation angle; 0c is the initial cohesion yield stress; a is the meridional

eccentricity with a default value of 0.1; and e is the deviatoric eccentricity which is

calculated as )sin3()sin3( φφ +−=e by default.

In this constitutive model, the Young’s Modulus, Poisson’s ratio, friction angle and

dilation angle need to be provided. The material properties for the shearing zone were

assumed based on the soil properties and verified by test results. The properties for all

of the materials in this finite element model are summarized in Table 8.1.

- 197 -


Actually in the pull-out test, the constrained dilatancy of the material in the shearing

zone was not constant throughout the test. After the peak pull-out resistance, the earth

pressures measured by earth pressure cells 1 to 4 were observed to have decreased

which indicated that the constrained dilatancy had decreased. In ABAQUS, the material

properties can be defined to depend on field variables and as a result, can be changed

during the analysis. In this model, the dilation angle of the material in the shearing zone

is defined to depend on the equivalent plastic strain using the subroutine of user defined

field. After the peak pull-out resistance, the dilation angle is gradually reduced to zero.

8.4 SIMULATION OF THE PULL-OUT TESTS

Using the model described above, simulations with applied overburden pressures of

40kPa, 80kPa, 120kPa, 200kPa and 300kPa were carried out to verify the model by

simulating the influence of the overburden pressure on the soil nail pull-out resistance.

The simulated results will be presented and discussed in the following paragraphs and

compared with the measured results in the laboratory pull-out tests.

8.4.1 Stress and strain rate contours

Figure 8.4 and Figure 8.5 show the vertical stress contours after drilling and after

pull-out, respectively, for a simulation with an applied overburden pressure of 120kPa.

In the stress contours, compressive stress is negative. From Figure 8.4, it can be

observed that the stresses are released after the elements at the location of the hole are

removed so that the vertical stresses in the elements above the hole decrease to zero. As

the distance from the surface of the hole increases, the vertical stress increases until it

reaches the applied overburden pressure of 120kPa. During pull-out of the nail, the

vertical stresses in the elements above the hole increase due to the constrained dilatancy

- 198 -


of the material. As a result, it can be observed that vertical stresses in the elements

above the nail become more even in vertical direction at the end of the pull-out (Figure

8.5). Figure 8.6 shows the maximum strain rate contour of the model after pull-out. The

maximum strain rates of the elements in the shearing zone are much larger than those in

other elements, which indicate that plastic flow had occurred in the shearing zone.

8.4.2 Variations of the vertical stress during the simulation

Figure 8.7 shows the vertical stress distribution before and after drilling along the path

shown in Figure 8.4. In Figures 8.7, 8.8 and 8.9, compressive stress is positive. Before

drilling, the vertical stress at the top surface of the model is equal to the applied

overburden pressure of 120kPa and it increases linearly as the depth increases due to

gravity effect. After drilling, the vertical stress in the elements just above the hole

decreases to zero. As the distance from the hole surface increases, the vertical stress first

increases quickly to about 90% of that before drilling at the distance of 0.15m and then

slowly increases to the applied overburden pressure at the top surface of the model. This

is similar to the measured earth pressures in the pull-out tests where the earth pressures

measured by earth pressure cells 5 and 6 just decreased in a small amount after drilling.

Figure 8.8 shows the distribution of the vertical stress and increased vertical stress at

peak pull-out resistance along the path shown in Figure 8.4. The distribution of the

vertical stress at peak pull-out resistance is more uniform compared with that after

drilling. The increased vertical stress is the maximum in the elements just above the nail

and decreases quickly to about 10kPa as the distance increases to 0.1m. As the distance

continues to increase, the increased vertical stress decreases slowly and become zero

when the distance is greater than 0.3m. In the pull-out tests, almost no change was

observed in the earth pressures measured by earth pressure cells 5 and 6 during pull-out.

In the finite element model, the increased vertical stress at peak pull-out resistance at

- 199 -


the level of earth pressure cells 5 and 6 (0.2m below the top surface) is just about 3kPa.

The agreement between the measured and simulated results is good.

Figure 8.9 shows the changes of average vertical stress at the locations of earth pressure

cells 1 to 4 at different stages of modeling under different applied overburden pressures.

Comparing with the results shown in Figure 6.4 (a), it can be seen that the typical

variations of the measured average earth pressure and the simulated average vertical

stress are similar. The values of the measured and simulated average earth pressure

(vertical stress) at each stage of testing (modeling) are also close to each other.

8.4.3 Influence of the overburden pressure

Figure 8.10 shows the simulated average pull-out shear stress-displacement curves

under different applied overburden pressures. The average pull-out shear stress was

calculated using the same method as in the laboratory pull-out tests. Only slight increase

in the peak average pull-out shear stress is observed as the applied overburden pressure

increases from 40kPa to 200kPa. The peak pull-out shear stress under the overburden

pressure of 300kPa is smaller than that under the overburden pressure of 200kPa. This is

consistent with the observation in the laboratory pull-out tests that the applied

overburden pressure did not significantly influence the pull-out resistance of the soil

nail for drill-and-grout nails. Figures 8.11 to 8.15 show the comparison between the

measured and simulated average pull-out shear stress-displacement curves under

different applied overburden pressures. The agreements between the measured and

simulated results are generally good. Before the peak pull-out resistance, the measured

average pull-out shear stress-displacement curves have a short linear part and then

become nonlinear. However, the simulated average pull-out shear stress linearly

increases to the peak pull-out resistance because the constitutive model used in the

- 200 -


simulation is a linear elastic-perfectly plastic model. In addition, the simulated peak

pull-out resistances are not perfectly the same as the measured ones for all the

overburden pressures. This is probably because that the test conditions cannot be

perfectly controlled due to some inevitable disturbance factors.

8.5 PARAMETRIC STUDIES

Parametric studies are performed on the influences of the dilation angle of the shearing

zone and the grouting pressure on the soil nail pull-out resistance. The results are

presented and discussed in the following paragraphs.

8.5.1 Influence of dilation angle

The curves shown in Figure 8.16 are the relationship between the simulated average

pull-out shear stress and pull-out displacement with different dilation angles under the

same applied overburden pressure of 120kPa. From the results, it can be observed that

the average pull-out shear stress increases linearly at the beginning of the pull-out

process, the larger the dilation angle, the larger the increasing rate. After a peak is

reached, the pull-out shear stress tends to keep constant till the end of the pull-out. The

peak pull-out shear stress with zero dilation angle is quite small which indicates that the

dilatancy of the soil plays an important role in development of the pull-out resistance.

Figure 8.17 shows the simulated peak pull-out shear stress vs. dilation angle under the

same applied overburden pressure of 120kPa. The peak pull-out shear stress is observed

to increase with the increase in dilation angle. The increasing rate decreases as the

dilation angle increases. Figure 8.18 shows the relationship between the simulated

vertical displacement on the top surface of the model and pull-out displacement with

different dilation angles under the applied overburden pressure of 120kPa. The curves

are similar to those in Figure 8.16. The vertical displacement first increases linearly

- 201 -


with the increase in pull-out displacement and then tends to keep constant at larger

pull-out displacement, the larger the dilation angle, the larger the maximum vertical

displacement.

This parametric study shows the theoretical influence of the dilation angle of the

shearing zone on the pull-out resistance of a drill-and-grout soil nail. Laboratory

pull-out tests are suggested to be carried out in soil with different dilation angles to find

out the actual influence of this parameter.

8.5.2 Influence of grouting pressure

The influence of grouting pressure is also investigated using the finite element method.

The simulation procedure is similar to that described in section 8.3.2. In the second

model, the calculation is accomplished in three steps. The first and second steps are the

same as those in the foregoing simulations. In the third step, a pressure load

representing the grouting pressure is applied to the surface of the hole which is formed

in the second step. After the calculation is completed, the stresses in all of the elements

are exported to be used as initial stresses for the third model. The third model includes

five steps. In the first step, the stresses obtained from the second model are imported

and overburden pressure and gravity loads are applied. The grouting pressure load

acting on the hole surface is removed and replaced by fixed boundary conditions. The

stresses generated by the grouting pressure are therefore locked in the elements

surrounding the hole. In the second step, the nail is activated without stress. The fixed

boundary conditions are removed in the third step and the locked-in stresses are released

and transmitted to the nail surface. The gravity of the nail is applied in the fourth step

and the nail is pulled out in the fifth step (Figure 8.19).

- 202 -


Figure 8.20 shows the minimum principal stress contour after application of grouting

pressure for a simulation with the applied overburden pressure of 80kPa and grouting

pressure of 200kPa (compressive stress is negative). The applied grouting pressure is

normal to the hole surface so that the minimum principal stress in areas surrounding the

hole should be normal to the hole surface. Therefore, in areas surrounding the hole, the

minimum principal stress represents the stress normal to the surface of the hole. From

Figure 8.20, it can be observed that the stress normal to the hole surface is approximate

to 200kPa in the elements next to the hole surface and decreases as distance increases.

Figure 8.21 shows the variation of the minimum principal stress in the elements next to

the hole surface throughout the simulation (compressive stress is negative). The

compressive stress decreases during excavation of the hole and then increases to a value

close to the grouting pressure after application of the grouting pressure. In the following

four steps, the compressive stress keeps constant. During pull-out, the compressive

stress increases due to constrained dilation of the material surrounding the nail.

Figures 8.22 to 8.24 show the simulated average pull-out shear stress – displacement

curves with different grouting pressures under the applied overburden pressures of

80kPa, 120kPa and 200kPa. For the overburden pressure of 80kPa, the simulation with

the grouting pressure of 400kPa cannot converge because that the grouting pressure is

too large compared with the overburden pressure. The increasing rates and peak pull-out

stresses of the curves are found to increase as the grouting pressure increases. Figure

8.25 shows the simulated peak pull-out shear stress vs. grouting pressure under different

applied overburden pressures. The peak pull-out shear stress is observed to significantly

increase with the increase in grouting pressure. The increase in peak pull-out shear

stress with the overburden pressure is insignificant.

- 203 -


The numerical simulation of pressure grouting only simulates the influence of the

locked-in grouting pressure acting on the nail surface. The simulation does not

reproduce the pressure grouting tests because that the grouting pressure dropped to zero

after grouting and the increase in pull-out resistance was due to change of properties for

soil and nail-soil interface. No directly measured data for the change of soil and nail soil

properties is available for numerical simulation. Therefore it is impossible to reproduce

the tests in the simulation. The FE modeling is probably over-simplified due to time

limit of the PhD study. More work is required to be carried out in the future to further

study the influence of grouting pressure on the soil nail pull-out resistance.

8.6 SUMMARY

A 3-D finite element model has been developed and verified by using pull-out test

results. Results of parametric studies on the influence of the grouting pressure and the

dilation angle of the shearing zone on the soil nail pull-out resistance are presented and

discussed. Based on the simulation and discussion, the following conclusions can be

drawn:

(a) The simulated results are in good agreements with the measured results from

the laboratory pull-out tests. This indicates that the use of continuum elements

for simulation of the nail-soil interface is applicable by using appropriate

material properties.

(b) The simulated peak pull-out resistance is significantly increased with the

increase in grouting pressure and dilation angle of the shearing zone. This

indicates the constrained dilatancy of the soil surrounding the nail and the

grouting pressure contribute a lot to the peak pull-out resistance.

- 204 -


Table 8.1 Material properties used in the finite element model

Density Young’s Modulus

Poisson’s ratio

Friction angle

Dilation angle Cohesion

Nail 2.0Mg/m3 16.37GPa 0.28 Plate 1.6GPa 0.25 Soil 1.668Mg/m3 14MPa 0.35 34° 11° 26kPa Shearing zone 1.668Mg/m3 14MPa 0.35 29° 10° 5kPa

1.0m

0.3m

0.81

m

21

3 Figure 8.1 –3-D mesh of the finite element model

- 205 -


Pressure

(a)

1

2 Pressure

(b) 3

2

Figure 8.2 – Load and boundary conditions of the FE model before pull-out – (a) front view (cross-section) and (b) side view (longitudinal section)

- 206 -


Pressure

Pull

3

2

Figure 8.3 – Load and boundary conditions of the model during pull-out (side view)

Path

Figure 8.4 – Vertical stress contour after drilling for a simulation with an applied overburden pressure (OP) of 120kPa (compressive stress is negative)

- 207 -


Figure 8.5 – Vertical stress contour after pull-out for a simulation with OP=120kPa (compressive stress is negative) Figure 8.6 – Maximum strain rate contour after pull-out for a simulation with OP=120kPa

- 208 -


0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0 50 100 150 200 250Vertical stress (kPa)

Dist

ance

from

nai

l sur

face

(m)_

Before drillingAfter drilling

Figure 8.7 – Vertical stress distribution before and after drilling along the path shown in Figure 8.4

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0 50 100 150 200 250Vertical stress (kPa)

Dist

ance

from

nai

l sur

face

(m)_

Vertical stress

Increased vertical stress

Figure 8.8 – Distribution of the vertical stress and increased vertical stress at peak pull-out resistance along the path shown in Figure 8.4

- 209 -


Changes of average vertical stress at different stages of modeling

0

50

100

150

200

250

300

350


Ave

rage

ver

tical

stre

ss (k

Pa)_

Before drillingAfter drillingBefore pull-outAt peak resistance Before drilling

At peak resistance

After drillingBefore pull-out

Figure 8.9 – Changes of average vertical stress at the locations of P-cells 1 to 4 at different stages of modeling under different applied overburden pressures


0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

40 kPa 80 kPa120 kPa 200 kPa300 kPa

Figure 8.10 – Simulated average pull-out shear stress vs. pull-out displacement under different applied overburden pressures

- 210 -



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

Test ABAQUS

OP=40kPa



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

Test ABAQUS

OP=80kPa


- 211 -



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

60

Test ABAQUS

OP=120kPa



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

Test ABAQUS

OP=200kPa


- 212 -



0

20

40

60

80

100

120


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

Test ABAQUS

OP=300kPa



0

40

80

120

160

200


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

ψ=0 ψ=4 ψ=8ψ=10 ψ=14 ψ=18ψ=26 ψ=29

Figure 8.16 – Simulated average pull-out shear stress vs. pull-out displacement with different dilation angles under OP=120kPa

- 213 -


0

20

40

60

80

100

120

140

160

180

0 5 10 15 20 25 30 35Dilation angle (Degree)

Peak

pul

l-out

shea

r stre

ss (k

Pa)_

Figure 8.17 – Simulated peak pull-out shear stress vs. dilation angle with OP=120kPa

Vertical displacement vs. pull-out displacement

0.00

0.40

0.80

1.20

1.60

2.00


Ver

tical

disp

lace

men

t (m

m) _

60

ψ=0 ψ=4 ψ=8ψ=10 ψ=14 ψ=18ψ=26 ψ=29

Figure 8.18 – Simulated vertical displacement on the top surface of the model vs. pull-out displacement with different dilation angles under OP=120kPa

- 214 -


Model 1

Initial stress

Model 2Step 1

Import Initial stress

Model 2Step 2

Drill hole

Model 2Step 3

Pressure grouting

Pressure

Model 3Step 1

Import stress

Fixed boundary

Model 3 Step 2

Add nail

Fixed boundary

Model 3Step 3Release

boundary

Model 3Step 4

Add nail gravity

Model 3 Step 5 Pull-out

Figure 8.19 – Procedure for simulating the pressure grouting Figure 8.20 – Minimum principal stress contour after application of grouting pressure for a simulation with OP=80kPa and grouting pressure (GP) of 200kPa (compressive stress is negative)

- 215 -


-400

-350

-300

-250

-200

-150

-100

-50

00 1 2 3 4 5 6 7 8 9

Step time

Min

imum

prin

cipa

l stre

ss (k

Pa) _

Figure 8.21 – Variation of the minimum principal stress in the elements next to the hole surface throughout the simulation with OP=80kPa and GP=200kPa (compressive stress is negative)


0

40

80

120

160

200


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

GP=0 GP=80 GP=130GP=200 GP=300 GP=400OP=80kPa


- 216 -



0

50

100

150

200

250


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

60

GP=0 GP=80 GP=130GP=180 GP=300 GP=400

OP=120kPa



0

40

80

120

160

200

240

280


Ave

rage

pul

l-out

shea

r stre

ss (k

Pa) _

0

GP=0 GP=80 GP=130GP=180 GP=300 GP=400

OP=200kPa


- 217 -


70

90

110

130

150

170

190

210

0 50 100 150 200 250 300 350 400 450Grouting pressure (kPa)

Peak

pul

l-out

shea

r stre

ss (k

Pa)_ OP=80 OP=120 OP=200

Figure 8.25 – Simulated peak pull-out shear stress vs. grouting pressure under different applied overburden pressures

- 218 -


CHAPTER 9: SUMMARY, CONCLUSIONS AND

SUGGESTIONS

9.1 SUMMARY

In this thesis, a literature review is made first on the soil nailing technique and its

applications. Types and characteristics of soil nails and application fields of the soil

nailing technique are summarized. The soil nailing mechanism and failure modes of soil

nail structures are presented. Current design methods and design guides are introduced.

Previous testing and numerical studies on the soil nailing technique and soil nail

structures are reviewed. Approaches available for estimating the soil nail pull-out

resistance and studies on factors influencing the pull-out behaviour and resistance are

reviewed in detail.

In the present study, elementary tests were conducted to simulate the pull-out behaviour

of a soil nail in soil slope. The tests were carried out using two copies of a newly

designed pull-out box in compacted completely decomposed granite (CDG) fill. The

internal dimensions of the pull-out box are 1.0m in length, 0.6m in width and 0.83m in

height. Totally 24 soil nail pull-out tests were carried out using the two pull-out boxes.

Among them, 20 tests were performed with the nails grouted under gravity head only

and 4 tests were carried out with the nails grouted with different applied grouting

pressures. The gravity grouting tests were conducted in soil at degrees of saturation of

38%, 50%, 75% and 98% (submerged) under overburden pressures of 40kPa, 80kPa,

120kPa, 200kPa and 300kPa respectively. The pressure grouting tests were carried out

in soil at 50% degree of saturation only, with grouting pressures of 80kPa and 130kPa

- 219 -

Chapter 9: Summary, conclusions and suggestions

under overburden pressure of 80kPa and 200kPa respectively. The test results were

presented and discussed.

For submerged tests, the test soil needed to be saturated before pull-out of the soil nail.

In order to obtain a higher degree of saturation, back water pressure needs to be applied

to the soil sample because of the large volume of the sample. Preventing leakage of

water is difficult during applying back water pressure to a soil nail pull-out box

especially at the location of the nail head because that the nail needs to be pulled out

through the nail head. Enlightened by the mechanism of a traditional triaxial apparatus,

a specially designed waterproof front cap was used to cover the nail head in submerged

pull-out tests. The waterproof front cap could prevent leakage of water under a relative

high (more than 100kPa) pressure and allowed the nail to be pulled out with very low

friction between the rubber O-ring and the steel rod for pull-out. During saturating the

soil and pull-out of the nail, no water leakage was observed and an average degree of

saturation of 98% was found to be achieved by checking the soil after pull-out.

The CDG soil was thoroughly remixed and compacted in the box and the extension

chamber in 9 layers (with the maximum thickness of 100mm) with 95% of the

maximum dry density (1.668Mg/m3). For each degree of saturation, the soil was mixed

with water to the target moisture content before compaction. The stress release due to

installation of the soil nail was simulated by drilling the hole after application of the

overburden pressure. The estimated peak pull-out load was divided into 5 loading

increments and applied stepwise. After the peak pull-out resistance was achieved, the

nail was continuously pulled out by displacement control using displacement rates of

1mm/min or about 0.3mm/min for tests in partially saturated and saturated soil

respectively.

- 220 -


The overburden pressure was applied by water pressure through a rubber diaphragm

fixed on the bottom surface of the top cover of the box and measured by pressure dial

gauge. An automatic volume change apparatus, which is normally used in the traditional

triaxial tests, was used to measure the volume change of the soil. Earth pressure cells

were imbedded in the test soil at different levels to estimate the distributions and

variations of stresses in the soil during testing. Soil moisture probes and miniature

porewater pressure transducers were installed to locations 25mm away from the nail

surface to measure the suction or porewater pressure in the soil surrounding the nail.

The axial strains at different locations of the soil nail were measured by means of strain

gauges adhered to the steel bar.

Numerical modeling was carried out using a three dimensional (3-D) finite element

model. The agreements between the measured and simulated results were good.

9.2 CONCLUSIONS

Based on the results of the pull-out tests and numerical simulation, the following

observations and conclusions can be obtained for soil nails installed in a compacted

completely decomposed granite (CDG) fill:

(a) Variation of the earth pressure throughout the test: After application of the

overburden pressure, the average earth pressure measured by pressure cells 1 to 4

was approximately equal to the applied overburden pressure. The average earth

pressure of pressure cells 1 to 4 substantially decreased to residual values after

drilling and did not recover much after installation of the soil nail. At the peak

pull-out resistance, the average earth pressures increased to certain values due to

- 221 -


the constrained dilatancy of the soil. This indicated that the normal stress acting

on the surface of the soil nail during pull-out, which contributed to the pull-out

resistance, was largely generated by constrained dilatancy of the soil.

(b) Average pull-out shear stress-displacement behaviour: The average pull-out

shear stress-displacement curves generally showed a peak and post-peak

displacement softening behaviour. The post-peak displacement softening was

more obvious in tests with soil at higher degrees of saturation. This is probably

related to the migration of shearing plane from the interface between the soil nail

surface and the surrounding soil to the interior of the soil when the soil is getting

wetter. For the tests performed in dryer soil, the failures occurred on the nail-soil

interface. The displacement softening was not the dominating behaviour of the

dilative soil subjected to shearing against a hard surface. For the tests performed

in wetter soil, the failures occurred within the soil. The pull-out

stress-displacement behaviour showed obvious displacement softening as

generally shown in the behaviour of dense dilative soil.

(c) Influence of overburden pressure on pull-out resistance: The peak pull-out

shear resistances for tests under different overburden pressures were a little

scattered and did not directly related to the applied overburden pressures. This

was consistent with the observations from Cartier and Gigan (1983) and

Clouterre (1991) by field pull-out tests. The pull-out resistances under lower

overburden pressure could be larger than those under higher overburden pressure.

The variation of the peak pull-out shear resistance under different overburden

pressures might have been caused by some inevitable disturbance factors, such as

- 222 -


misalignment of the pull-out force, disturbance to the hole surface due to drilling,

and etc.

(d) Influence of degree of saturation on pull-out resistance: The peak pull-out

shear resistance varied with different degrees of saturation of the soil, with higher

resistances at the degrees of saturation of 50% and 75%. In both dryer (38%) and

wetter (submerged) soil, the pull-out shear resistance was lower. It is believed

that if the soil is dry, water is more readily to be absorbed by the soil due to a

higher suction power. This may cause more contraction of the cement grout thus

reducing the bond strength between the nail surface and the surrounding soil. As

in the submerged tests, the decrease of the apparent cohesion of the fill could be

one of the reasons for the decrease of pull-out resistance. The decrease of

dilatancy of the soil could be another reason.

(e) Effect of pressure grouting: The grouting pressure inside the hole increased

quickly, but then dropped rapidly due to hardening of the cement grout and

shrinkage of grout volume because of water seepage from the cement grout to the

surrounding soil (accompanied by dissipation of the excess water pressure inside

the grout and absorbing of water in the grout by the dryer surrounding soil). This

indicated that the grouting pressure did not permanently increase the normal

stress acting on the nail-soil interface.

(f) Influence of grouting pressure on pull-out resistance: Even though the

grouting pressure just temporarily increased the normal stress acting on the

nail-soil interface, the pull-out resistance increased with the increase in grouting

pressure. The increase in pull-out resistance was probably due to the

- 223 -


recompaction of the soil around the hole, the infiltration of the cement grout into

the soil, the increase in nail surface roughness and etc.

(g) Numerical modeling: Numerical modeling results using a 3-D finite element

model are in good agreement with measured results. This indicates that using a

thin layer of continuum elements with Mohr-Coulomb material to simulate the

nail-soil failure surface is applicable. Parametric study shows that the pull-out

resistance generally increases with the increase in grouting pressure and dilation

angle of the material in the shearing zone.

9.3 RECOMMENDATIONS AND SUGGESTIONS

(a) The pull-out tests in this project were carried out in CDG soil at only four

different degrees of saturation and could not show a clear relationship between

the soil nail pull-out resistance and degree of saturation of the soil. More tests

are recommended to be carried out in soil at more different degrees of saturation

in order to find out the relationship.

(b) The earth pressure cells close to the nail were imbedded at locations about

40mm away from the nail surface to avoid damage of the pressure cells. The

measured earth pressures cannot represent the normal stress acting on the

nail-soil interface. Direct measurement of the normal stress on the nail-soil

interface may be to be developed to study the influence of the normal stress on

the pull-out resistance.

(c) The bond strength between the cement grout of the soil nail and the surrounding

soil may depend on the curing time of the cement grout. Tests with soil nails of

- 224 -


different curing time are suggested to be carried out to study the influence of

curing time on the pull-out resistance.

(d) Only four pressure grouting pull-out tests were carried out in the present work.

More pressure grouting tests are recommended to be conducted with different

grouting and overburden pressures in soil at different degrees of saturation.

Especially, pressure grouting tests should be carried out in submerged

conditions to see whether or not the grouting pressure can still increase the

pull-out resistance when the soil is saturated.

(e) In a soil nailed slope, the nail in the passive zone is normally subjected to a

combination of shear force, tensile force and bending moment when a failure

surface is formed. Pull-out tests is recommended to be carried out considering

the influence of shear force and bending moment.

(f) Pull-out tests are recommended to be carried out on nails with different

diameters and surface roughness to study the influences of these factors on the

pull-out resistance.

(g) The pull-out test results showed that the constrained dilatancy of the soil

contributed a lot to the pull-out resistance. Analytical solution of the pull-out

resistance considering the influence of dilation angle of the soil is recommended

to be developed and used for design and numerical modeling.

- 225 -

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