Interaction Properties of Geogrids in Reinforced Soil

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LSU Historical Dissertations and Theses Graduate School

1990

Interaction Properties of Geogrids in ReinforcedSoil Walls, Testing and Analysis.Khalid A. FarragLouisiana State University and Agricultural & Mechanical College

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O rd er N um ber 9123187

Interaction properties o f geogrids in reinforced-soil w alls, testin g and analysis

Faxrag, K h a lid A ., P h .D .

The Louisiana State University and Agricultural and Mechanical Col., 1990

U MI300 N. Zeeb Rd.Ann Arbor, MI 48106



INTERACTION PROPERTIES OF GEOGRIDS IN REINFORCED-SOIL WALLS,TESTING AND ANALYSIS.

A Dissertation

Submitted to the Graduate Faculty of the Louisiana State University and

Agricultural and Mechanical College in partial fulfillment of the degree of

Doctor of Philosophy

in

The Department of Civil Engineering

byKhalid A. Farrag

B.S. in C.E., Cairo University, 1977 M.S. in C.E., Louisiana State University, 1986

Dec. 1990


ACKNOWLEDGMENT

The author wishes to express his sincere thanks to his major advisor Dr.

Ilan Juran for his guidance and supervision during the period in which this study

was carried out, and to Dr. Y. Acar, the committee Chairman, for his advice

through the last portion of this study. My appreciation also extends to the

other members of the committee for their suggestions.

The author wishes to thank the Board of Regent, Louisiana

Transportation Research Center (LTRC), and Conwed Plastics for their financial

support. The suggestions from the staff of LTRC and Conwed Plastics during

testing are also highly appreciated.

Finally, sincere appreciation is forwarded to my wife Elisabete for her

help and support throughout my study.

ii


TABLE OF CONTENTS

Page

ACKNOWLEDGMENT. ii

TABLE OF CONTENTS. iii

LIST OF TABLES. vii

LIST OF FIGURES. xiii

ABSTRACT. xv

CHAPTER 1 INTRODUCTION 1

CHAPTER 2 REVIEW OF THE EXISTING GEOSYNTHETIC TESTING 7PROCEDURES.

2.1 IN-SOIL MECHANICAL PROPERTIES OF GEOSYNTHETICS. 7

2.2 SOIL-GEOSYNTHETIC INTERFACE PROPERTIES. 16

2.3 CONCLUSIONS. 29

CHAPTER 3 EQUIPMENT DESIGN AND INSTRUMENTATION. 32

3.1 DESIGN CONSIDERATIONS. 32

(i) Loading Scheme. 32

(ii) Rigid Wall Boundary Effects. 33

(iii) Soil Placement and Compaction. 33

(iv) The Specimen Clamping Mechanism. 34

(v) Instrumentation. 34

3.2 EQUIPMENT DESCRIPTION. 35

(i) The Pull-out and Shear Box Details. 36

(ii) The Hydraulic Loading System. 43


Ease

(iii) Sand Handling Facility. 44

3.3 INSTRUMENTATION AND DATA ACQUISITION. 48

(i) Load and Pressure Measurements. 49

(ii) Displacement Measurements. 50

(iii) Displacement-rate Measurements. 51

(iv) Data Acquisition. 53

CHAPTER 4 TESTING PROCEDURE AND DATA INTERPRETATION. 61

4.1 INTRODUCTION. 61

4.2 SOIL COMPACTION AND SAMPLE PREPARATION. 62

(i) Soil Placement. 63

(ii) Compaction Control. 64

(iii) Reinforcement Preparation. 66

4.3 DISPLACEMENT-RATE CONTROLLED TESTS. 75

(i) Unconfined Extension Tests. 75

(ii) Pull-out Tests. 75

4 .4 DIRECT SHEAR TESTS. 84

4.5 LOAD-CONTROLLED TESTS. 86

(i) Unconfined Load-Controlled Extension Tests. 87

(ii) Stepped Load-Controlled Pull-out Tests. 88

(iii) Long Term Pull-out Tests. 89

iv


Page

CHAPTER 5 EFFECT OF TESTING PARAMETERS ON THE 96INTERACTION MECHANISM.

5.1 INTRODUCTION 96

5.2 EFFECT OF TESTING PARAMETERS ON INTERACTION 101RESPONSE.

(i) Reinforcement Type. 101

(ii) Displacement Rate. 103

(iii) Soil Compaction and Relative Density. 110

(iv) Confining Pressure. 114

(v) Sleeve Length. 120

(iv) Thickness of Soil. 126

CHAPTER 6 MODELING LOAD TRANSFER MECHANISM. 129


6.2 MODELING LOAD TRANSFER IN PULL-OUT TESTS. 138

6.3 ANALYSIS OD LOAD-CONTROLLED PULL-OUT TESTS. 152

CHAPTER 7 STABILITY ANALYSIS OF REINFORCED SOIL 159STRUCTURES.


7.2 MODELING SOIL-GEOSYNTHETIC SOIL WALLS. 163

7.3 ANALYSIS OF BEHAVIOR OF GEOSYNTHETIC SOIL WALLS. 171

7.4 ANALYSIS OF MODEL WALL RESULTS. 178

(i) Analysis of Laboratory Model Tests. 178

(ii) Analysis of Full-Scale Model Walls. 183

v


Page

(iii) Analysis of Reinforced Embankments. 190

CHAPTER 8 CONCLUSIONS. 196

REFERENCES. 200

APPENDIX A EQUIPMENT DETAILS AND SPECIFICATIONS. 208

APPENDIX B COMPACTION CONTROL AND CALIBRATION. 218

APPENDIX C PULL-OUT TEST RESULTS. 224

APPENDIX D FORMULATIONS OF DESIGN MODEL EQUATIONS. 259

VITA. 267

vi


LIST OF TABLES

Table Page

2.1 Comparison of Geosynthetic Direct Shear Tests. 19

2.2 Comparison of Geosynthetic Pull-out Tests. 24

5.1 List of Pull-out Tests. 98

5.2 Comparison Between Strain Rates for Different 106Pull-out Tests.

7.1 Material Properties and Geometrical Parameters 179of Test Walls.

7.2 Material Properties and Geometrical Parameters 184of Model Walls.

7.3 Properties of Reinforced Embankments. 190


LIST OF FIGURES

Fiaure Paqe

2.1 Unconfined and Confined Extension Tests on Geotextiles. 11

2.2 Modified Shear Box for Confined Extension Tests. 12

2.3 Confined Extension Testing Device on Geotextiles. 12

2.4 Confined Extension Testing Apparatus. 13

2.5 Confined Extension Testing Apparatus. 14

2.6 'Zero Span' Confined Extension Testing Apparatus. 15

2.7 Triaxial Cell with Reinforced Soil Sample. 15

2.8 Soil-Geosynthetic Direct Shear Test Devices. 18

2.9 Pull-out Testing Device. 22

2.10 Pull-out Testing Device. 22

2.11 Geosynthetic Pull-out Box. 23

2.12 Displacement Monitoring Along the Reinforcement in Pull-out. 23

2.13 Comparison of the Efficiency Factor from Direct Shear and Pull-out Tests.

31

3.1 Side View of the Pull-out Box. 38

3.2 Front View of the Pull-out Box and Loading Frame. 39

3.3 Longitudinal Cross Section of the Pull-out Box. 40

3.4 Cross-Section of the Pull-out Box. 40

3.5 View of the Large Direct Shear Box. 41

3.6 Cross-Section of the Direct Shear Box. 42

viii


Figure Pb99.

3.7 View of the Loading Frame. 45

3.8 Details of the Hydraulic Loading System. 46

3.9 View of the Sand Handling Facility. 47

3.10 Data Acquisition Supporting Instruments. 55

3.11 Locations of the Earth Pressure Cells. 56

3.12 Detail of the 'Tell Tail' Rear Table. 56

3.13 Displacement Instrumentation of the Geogrid. 57

3.14 Schematic View of the Data Acquisition System. 58

3.15 View of the Data Control System. 59

3.16 Connection Scheme of the Screw Terminal Board. 60

4.1 Grain Size Distribution of the Blasting Sand. 68

4.2 Direct Shear Test Results on Sand. 69

4.3 Sand Placement in the Pull-out Box. 70

4 .4 Soil Density Measurement in the Pull-out Box. 71

4.5 Calibration of Sand Density with Compaction. 72

4.6 Placement of the Geogrid in the Pull-out Box. 73

4.7 Instrumentation of Geogrid Specimens in the Pull-out Box. 74

4.8 Unconfined Extension Test Results on Geogrid. 79

4.9 Pull-out Test Results on Geogrid'Tensar SR2'. 80

4.10 Pull-out Test Results on Geogrid 'Tensar SR2'. 80

ix


Figure Page

4.11 Pull-out Test Results on'Conwed'Geogrid. 81

4.12 Measurement of Front Displacement in Pull-out Tests. 81

4.13 Time-Nodal Displacement Relationship for 'Tensar' Geogrid. 82

4.14 Time-Nodal Displacement Relationship for 'Conwed' Geogrid. 82

4.15 Normalized Displacement Along the 'Tensar' Geogrid. 83

4.16 Normalized Displacement Along the 'Conwed' Geogrid. 83

4.17 Effect of Reinforcement on the Frictional Resistance 85in Direct Shear Test.

4.18 Illustration of Creep Concept. 90

4.19 Determination of the Critical Creep Load. 90

4.20 Concept of Strain Superposition in Stepped Loading Tests. 91

4.21 Unconfined Load Controlled Tests on Geogrid Tensar. 92

4.22 Stepped Load-Controlled Test on Geogrid'Conwed'. 93

4.23 Stepped Load-Controlled Test on Geogrid 'Conwed'. 93

4.24 Log (Displ. rate) versus Log (Time) in Creep Tests. 94

4.25 Determination of the Critical Creep Load. 94

4.26 Results of Creep Test on Geogrid 'Conwed'. 95

4.27 Results of Creep Test on Geogrid'Conwed'. 95

5.1 Load Transfer Mechanism in Geogrids. 102

5.2 Pull-out Tests on Geogrids With and Without Transversal Ribs. 102

5.3 Effect of Displacement-Rate on Pull-out Response. 107

x


Figure Pggg

5.4 Effect of Displacement-Rate on Nodal Displacement. 107

5.5 Effect of Displacement-Rate on Displacement Distribution. 108

5.6 Effect of Displacement-Rate on Pull-out Resistance. 109

5.7 Effect of Displacement-Rate on Interface Stiffness Modulus. 109

5.8 Effect of Soil Density on Pull-out Resistance of 'Tensar'. 112

5.9 Effect of Soil Density on Pull-out Resistance of 'Tensar'. 112

5.10 Effect of Soil Density on Displacement Distribution 113Along Geogrid

5.11 Effect of Confining Pressure on Pull-out Resistance of 'Tensar'. 117

5.12 Effect of Confining Pressure on Extension Behavior of 'Tensar'. 117

5.13 Effect of Confining Pressure on Pull-out Resistance of 118'Conwed' Geogrid.

5.14 Effect of Confining Pressure on the Displacement Distribution 119Along 'Tensar' Geogrid.

5.15 Effect of Confining Pressure on the Displacement Distribution 119Along 'Conwed' Geogrid.

5.16 Effect of Sleeve Length on Pull-out Response of'Tensar'. 122

5.17 Development of Earth Pressure on the Front Wall in Case of 123No-Sleeve.

5.18 Development of Earth Pressure on the Front Wall in Case of 1238 in. Sleeve.

5.19 Development of Earth Pressure on the Front Wall in Case of 12412 in. Sleeve.

5.20 Effect of Sleeve Length on Lateral Earth Pressure. 125

xi


Figure ESS£

5.21 Effect of Sleeve Length on Normalized Displacement. 125

5.22 Effect of Soil Thickness on Pull-out Response. 128

5.23 Effect of Soil Thickness on the Normalized Displacement. 128

6.1 Deformations Along The Geotextile. 136

6.2 Distribution of Pull-out Resistance Along the Geotextile. 136

6.3 Applied Load versus Front and Rear Displacement. 137

6.4 Numerical Simulation of Pull-out Test on Woven Strips. 137

6.5 Displacement Distribution Along Geogrid'Tensar'. 142

6.6 Displacement Distribution Along Geogrid 'Tensar'. 143

6.7 Displacement Distribution Along Geogrid 'Tensar'. 144

6.8 Numerical Analysis of Pull-out Tests. 145

6.9 Confined and Unconfined Stress-Strain Relationship for 146Geogrid 'Tensar'.

6.10 Confined and Unconfined Stress-Strain Relationship for 146Geogrid 'Conwed'.

6.11 Interface Shear Stress-Displacement Relationship for 147Geogrid 'Tensar'.

6.12 Interface Shear Stress-Displacement Relationship for 148Geogrid 'Tensar'.

6.13 Interface Shear Stress-Displacement Relationship for 149Geogrid 'Conwed'.

6.14 Analytical and Experimental Interface Shear Stress-Displ. 150Relationship for Geogrid 'Tensar'.


Figure Page

6.15 Analytical and Experimental Displacement Distribution Along 151the Geogrid Length.

6.16 Analytical Pull-out Load Distribution Along The Geogrid. 151

6.17 Nodal Displacements at Load Controlled Pull-out Test. 155

6.18 Nodal Displacements at Load Controlled Pull-out Test. 155

6.19 Confined & Unconfined Creep Strain for Geogrid 'Conwed'. 156

6.20 Confined & Unconfined Creep Strain for Geogrid 'Conwed'. 156

6.21 Strain-rate versus Time Relationship. 157

6.22 Strain-rate versus Stress Relationship. 157

6.23 Strain-rate versus Time for Geogrid 'Conwed'. 158

6.24 Strain-rate versus Stress for Geogrid'Conwed'. 158

7.1 Schematic Diagram of Failure Mechanisms in Reinforced Soil. 168

7.2 Analogy between Behavior of Reinforced Soils in Retaining 169Wall and in Direct Shear Test.

7.3 Elasto-Plastic Strain Hardening of Soil. 169

7.4 Modeling Behavior of Reinforcing Soils. 170

7.5 Flow Chart of Computational Scheme. 175

7.6 Effect of Reinforcement Stiffness on Maximum Tension Forces. 176

7.7 Effect of Soil Dilatancy on Maximum Tension Forces. 176

7.8 Effect of Reinforcement Stiffness on the Inclination 177of Failure Surface.

7.9 Observations on Inclination of Failure Surface for Different 177Reinforcements.


Figure Page

7.10 Geometry of Test Walls and Failure Surfaces. 180

7.11 'Confined' and 'Unconfined' Extension Tests on Non-Woven Geotextiles.

180

7.12 Measured and Predicted Tension Forces in Wall No. 1. 181


7.14 Geometry and Soil Properties of Model Walls. 185




7.18 State of Stress in the Reinforcement with Depth. 189

7.19 Measured and Predicted Tension Forces in Embankment 1. 191




7.23 Measured and Predicted Failure Surface in Embankment 1. 195

xiv


ABSTRACT

The considerable development of the geosynthetics and their increasing

use in soil reinforcement in the last decade made it necessary to develop

methods of measuring the interaction properties and modeling load transfer in

reinforced soil structures. The current testing procedures and interpretation

schemes demonstrate significant limitations related to the equipment design,

testing methodology and data interpretation. In order to address these

limitations, a pull-out box and direct shear box were constructed and

instrumented. The design of these boxes overcomes most of theses limitations

and provides the capability of conducting displacement-rate controlled tests for

the evaluation of short-term behavior of the soil-reinforcement system, and

load-controlled tests to evaluate the confined time-dependent performance of

the geosynthetics. The accuracy of the testing facility was evaluated through

comparison of test results with those provided by the manufacturers. The

reproducibility of test results were evaluated through repetitive tests performed

on the geogrids under the same testing parameters. A parametric study on the

geogrids under different testing conditions (i.e pull-out displacement-rate,

confining pressures, soil compaction, soil density and boundary conditions) was

conducted to evaluate the equipment sensitivity to the variations in testing

conditions and to provide a reliable data base for the evaluation of the effect

of these parameters on the soil-geogrid interface properties.

xv


A data analysis procedure and interpretation methodology were

established to determine the interface properties and the confined material

characteristics of the geogrid reinforcement. The applicability of the load-

transfer model was evaluated through comparison between the interface

parameters evaluated from pull-out tests with those measured directly from the

tests performed in the large direct shear box.

The interface properties and confined material characteristics of

the geogrid reinforcement were utilized in a design model for the stability

analysis of reinforced soil structures. This model is implemented in a computer

program to calculate the forces in the reinforcements and the predicted failure

planes. Design charts are presented where the maximum tension force at each

reinforcement level and the inclination of failure surface can be predicted for

different soil characteristics and reinforcement extensibilities. In order to

evaluate the design assumptions considered in this model, the predicted tension

forces and inclinations of failure surfaces were compared with those measured

in laboratory model tests, full-scale model walls, and embankments of different

soil data and reinforcements.

xvi


CHAPTER 1

INTRODUCTION

The considerable growth in using various materials to reinforce soil made

it necessary to develop methods of measuring the interaction properties and

modeling the load transfer mechanism at the soil-reinforcement interface. The

selection criteria for a specific material in reinforcement of the soil is governed

by the pull-out performance of the reinforcement, its tensile strength, and its

time-dependent behavior when placed in the soil. Two different types of

materials are used as tension resisting elements in reinforcing the slopes and

embankments; namely,

i) inextensible reinforcement (such as metal strips and welded wire meshes),

ii) extensible geosynthetic reinforcement (such as woven and non-woven

geotextiles and geogrids).

The use of inextensible materials in soil reinforcement have been investigated.

Studies led to development of well-established testing procedures for evaluation

of design parameters for soils reinforced with metal strips (Schlosser and Elias,

1978; Forsyth, 1978; Elias, 1979; and Schlosser et al., 1983). However, the

effect of corrosion on the long term durability of metal inclusions limited their

application and led to the use of polymer extensible materials. These extensible

materials have different constitutive properties and structures that influence

their behavior and consequently the testing methodology and interpretation.

1


Inextensible reinforcement moves as a rigid member in the soil with a

constant shear stress distribution along its length. While extensible

reinforcement exhibits more complex soil interaction mechanism. Non-uniform

shear stress and displacement distributions are developed along the soil-

inclusion plane with a large portion of the shear being mobilized at the loading

application point (Juran et al., 1988; Juran and Chen, 1988; Christopher,

1976; and Tzong and Cheng-Kuang, 1987). This non-uniform shear stress

distribution raises a basic difficulty in evaluating the scale effect for a reliable

interpolation of laboratory pull-out test results to field conditions.

In addition, all polymer materials exhibit time-dependent behavior and

stress relaxation which affect the long term stability of the reinforced

structures. Therefore, the interpretation methods established for inextensible

inclusions can not apply to the development of appropriate design procedure

for the extensible inclusions.

A wide variety of geotextiles and geogrids are now available for civil

engineering application (Bonaparte et al., 1987). In selecting a specific

geotextile or geogrid for reinforcement of embankments and slopes, the

following aspects of performance should be considered:

i) stress-strain relationship and creep behavior of the reinforcement when

placed in the soil,

ii) pull-out performance of the reinforcement and its load transfer mechanism.


The majority of stress-strain and creep property tests of geotextiles and

geogrids are currently performed on unconfined samples (Shrestha and Bell,

1982a-b; Andrawes et al., 1986; Richards and Scott, 1986; Rowe and Ho,

1986; and Myles, 1987). Consequently, the measured properties differ from

those obtained under confined conditions and can not be used to determine the

appropriate design parameters for the soil-reinforcement systems. Moreover,

no standard testing procedure for the properties of the confined reinforcement

currently exists. The large number of factors that affect the interface properties

of the confined reinforcement raises major difficulties in comparing test results.

Knochenmus (1987) and Juran et al. (1988) showed a wide scatter in the

available confined pull-out test results. These differences in results are partially

due to the use of different pull-out devices, the associated boundary effects,

testing procedures, and soil placement and compaction schemes.

In order to develop a methodology for evaluation of the in-soil mechanical

characteristics and interface properties of geosynthetic reinforcement, it is

necessary to establish reliable testing equipment, procedure, and appropriate

interpretation scheme. The specific objectives of this research are:

(1) to design, construct and calibrate a large pull-out box and a large direct

shear box which will provide the capability of evaluating the performance

of different types of geogrids and geotextiles,


4

(2) to develop a reliable testing procedure and data interpretation method for

evaluation of the short term and long term pull-out performance of

geosynthetic reinforcements,

(3) to utilize the measured interface properties and the in-soil material

characteristics in a design model in order to evaluate the stability of

reinforced soil structures.

The scope of this research focuses on testing and performance

evaluation of the geogrid reinforcements in dense granular soils. It involves five

tasks:

(1) Review of the Available testing Procedures: A comprehensive review of

the available testing equipment, procedures, and data interpretation methods

is conducted in order to evaluate the limitations of the current state of practice

and to provide guidelines for the design of the testing facility.

(2) Design of a Testing Facility: A testing facility is designed, constructed,

and instrumented for evaluation of both short term and long term geogrid

performance. The facility is designed to provide the capability of conducting

two basic testing modes: (a) displacement-rate controlled and (b) load

controlled modes. In the displacement-rate controlled mode, the geogrid is

subjected to a constant pull-out displacement rate during the test and pull-out

load is recorded. This testing procedure (which is most commonly used)

provides the interface parameters related to the short term performance of the


reinforcement such as peak and residual pull-out resistance and the level of

front displacement at the peak pull-out load. In the load controlled mode, pull-

out loads are applied incrementally to the inclusion and maintained constant

during a specified period. The displacements along the inclusion are recorded

and data interpretation yields time-dependent response parameters related to

the long term performance of the inclusion.

(3) Performance Evaluation Study: A pull-out and direct shear testing

programs were implemented in order to evaluate the performance of the facility

(i.e. the reproducibility of the results and accuracy of the monitoring system),

to assess the effect of the rigid boundaries on test results, and to provide a

data base for development of testing procedures (e.g. soil compaction, loading

mode and displacement rate).

(4) Development of Data Analysis Procedure: A data interpretation scheme

is developed in order to determine the interface properties and confined material

characteristics of geogrid reinforcements.

(5) Soil-reinforcement Design Model: The soil-reinforcement interface

properties are implemented in a computer program in order to analyze the state

of stresses in the reinforcement and to evaluate the stability of the reinforced

soil structures.

The detailed literature review of the existing testing equipment is

presented in Chapter 2. The design and instrumentation of both the pull-out and


direct shear boxes are presented in Chapter 3. Sample preparation, testing

procedure and performance evaluation of test results are presented in Chapter

4. A parametric study to assess the effect of different testing parameters on

the in-soil performance of the geogrids is presented Chapter 5. Chapter 6

provides an interpretation scheme and data analysis procedure in order to

determine the soil-geogrid interface parameters needed to evaluate the stability

of reinforced soil structures. Chapter 7 contains a design model for the stability

analysis of reinforced soil structures. Chapter 8 contains conclusions on the

pull-out and direct shear testing procedures, the effect of testing parameters

on soil-geogrid interaction mechanism, and a discussion of the data analysis

procedure and the stability of reinforced-soil structures.


CHAPTER 2

REVIEW OF THE EXISTING GEOSYNTHETIC TESTING PROCEDURES

The in-soil mechanical characteristics and interface properties of

geosynthetics have been experimentally modelled by monitoring:

i) the in-soil mechanical properties of the geosynthetic (i.e. its confined

stress strain relationship and creep behavior),

ii) the soil-geosynthetic interface properties (i.e. shear stress-strain

relationship and pull-out resistance).

Research on these parameters has been conducted using various equipment and

testing procedures which made it difficult to consistently compare the

performance of different geosynthetic specimens. An evaluation of these

equipment and testing procedures is presented herein.

2.1) IN-SOIL MECHANICAL PROPERTIES OF GEOSYNTHETICS:

The unconfined stress-strain properties of geotextiles can be determined

by testing the geotextile specimen in a wide width strip test according to the

procedures presented in ASTM D4595-86. However, when the geosynthetic

materials are embedded in the soil, their stress-strain properties are significantly

affected by soil confinement (McGown et al., 1982; Knochenmus, 1989; and

Juran et al., 1988).

7


The results of confined extension tests performed by Siel et al. (1987)

and McGown et al. (1982) show that the confining pressure increases the

tensile strength and deformation modulus of the geotextile material. These

results are presented, respectively, in Figures (2.1-a and b). The figures

demonstrate the effect of confining pressures on the deformation modulus. The

effect is more predominant on geogrids which have transversal elements. In

such geogrids, lateral earth resistance on the transversal elements restricts

elongation and hence, increases strength and deformation modulus..

Currently, no standard testing procedure or apparatus exists for

measuring the "in-soil" stress-strain material properties. Several investigators

have determined these properties by testing geotextiles in modified direct shear

boxes. In these boxes, the rear end of the specimen is clamped to the back of

the box; while the front of the specimen is subjected to an extension force. Siel

et al. (1987) used a modified shear box of 6.5 inch long, 4.75 inch wide and

2.5 inch deep (Figure 2.2) to study the stress-strain properties of non-woven

geotextiles in compacted sand. While, El-Fermaoui and Nowatzki (1982) used

a modified shear box of 2.5 inch long, 2.5 inch wide and 1.5 inch deep to

obtain the confined extension properties of woven and non-woven geotextiles.

The disadvantage of using small shear boxes with dimensions similar to

the conventional ones is the boundary effect of the box walls on the soil-

geotextile interaction. Moreover, such small boxes can not be used to evaluate

the elongation of the large representative samples of geogrids.


Another major shortcoming in using such devices is that confined

extension test produces a combined effect of shear stress and extension along

the specimen; thereby inducing a non-uniform tension force distribution along

the geotextile (Knochenmus, 1989). Additional instrumentations to measure

displacement distribution along the specimen and the load at its rear end are

necessary to de-couple the shear-extension effect.

Leshchinsky and Field (1987) used a direct shear box modified to contain

a metal platform in the lower box (Figure 2.3) to ensure that the geotextile

remains exactly at the shear plan. They measured the tension forces in both

ends of the geotextile specimen to estimate displacement and friction force

distribution along its length.

McGown et al. (1982) developed a custom-built apparatus (Figure 2.4)

to evaluate the effect of confining pressure and specimens size on the confined

extension properties of geotextiles. Their apparatus consists of two air pressure

diaphragms which are placed on each side of the geotextiles. A soil layer can

be compacted between the diaphragm and the geotextile. Their results, shown

in Figure (2.1 -b), demonstrate the effect of confinement and sample dimensions

on the confined properties of the geotextiles.

Knochenmus (1989) developed a confined extension testing device

(Figure 2.5) in which extension loads on both sides of the specimen and the

developed pore pressures in the soil sample can be monitored.


10

Attempts have also been made to determine the confined stress-strain

properties of the geotextiles using other devices. Christopher et al. (1986)

proposed a 'zero span' confined tension test in which the specimen is confined

by means of pressure controlled metal clamps. Their apparatus is shown in

Figure (2.6). Surface treatments were used on the clamp faces to simulate

granular soil conditions. Several shortcomings of this apparatus are related to

the difficulty in simulating the frictional conditions of the wide range of granular

and soft soils on the clamp surfaces. Moreover, the device does not account

for many parameters influencing the confined extension properties such as the

development of pore pressures in saturated fine soils, soil dilatancy and soil

particle interlocking on the specimen surface. Triaxial tests have also been

carried out to investigate the stress-strain properties and time-dependent

behavior of samples reinforced with horizontal disks of fabrics (Holtz et al.,

1982; and Broms, 1977). Figure (2.7) shows the reinforced soil sample in the

triaxial test. In this test, the strains in the fabric are difficult to be monitored

and they are associated with different boundary conditions than those

encountered in the field.

Although the "in-soil" stress-strain properties can be determined in

confined extension tests, the development of an appropriate load transfer

mechanism raises the need for other tests to determine the geogrid interface

properties (i.e. the shear stress-strain and pull-out resistance).


6 0 0 Ottawa #30 sand Monterey # 0 /3 0 sand

500Overburden Pressure

114.0 KPa400z

300T>DO

In-Air (0 KPa)200

100

0 2 4 6 8

Displacement (m m )(a) After Siel et al., 1987.

Unconfined In Isolation200*100 mm SPECIMEN^- Confined In Soil / s

confinig „ /stress (kN/nr/) .io o / / a

/ size of r test

specimens mml

500x100 400x100 300x100 200x100 100x100

0 10 20 30 40AXIAL STRAIN (*/.)

(b) After McGown et al., 1982.

Figure 2.1 Unconfined and Confined Extension Tests on Geotextiles.


12

Geotextile.

- 7 ^

N Top Loading CapA___/ - . Fixed______________ £ Clamp

S o i l \

P<-

MovableClamp

7i • • \ . • » . • • ,

• • • ' . V . - ' - . - - . - . - - : - .Wit• • • • • • • ■ • • I. . • . T 7 * . •. • 1

[ 7

S o i l V. - V■>P

Side View

Figure 2.2 Modified Shear Box for Confined Extension Tests. (After Siel et al., 1987)

r i n t e cl#uups r e i n f o r c e d e nd

co l owe r p l a c f o r a

OULUL C o n f i n i n g p r e s s u r e

\ S p a c e r \

A1u n i n u n p l a c e

Load a p p l i e d by Sc rew J a c k

P l a c e s on r e L n f o r c e a e n c

a o u n c I n g e o c e x c l l e l o a d

c e l l wl c h a p i n .

Figure 2.3 Confined Extension Testing Device on Geotextiles. (After Leshchinsky and Field, 1987)


■a- O

7T7 "OT

Figure 2.4 Confined Extension Testing Apparatus. (After McGown et al., 1982)


14

Soli

- - Bottom Clomp

— ~ ~ Boss Plolo

(T) Pon Picitura © Cell Praaaura (3) Back Pinturi

Figure 2.5 Confined Extension Testing Apparatus.(After Knochenmus, 1989}


Load

loveable TopSwivel

Clamp

fabric

AirLines

Figure 2.6 'Zero Span' Confined Extension Apparatus.(After Christopher et al., 1986)

Load

— P is to n

S / / / / —Top

— Cel l

G e o t e x t i l e

Flu id

Oase

Figure 2.7 Triaxial Cell with Reinforced Soil Sample, (after Broms, 1977)


16

2.2) SOIL-GEOSYNTHETIC INTERFACE PROPERTIES:

The shear stress-strain relationship developed at the soil-reinforcement

interface can be tested in both direct shear and pull-out boxes. In the direct

shear box, tests are usually conducted in accordance with the conventional

procedure of tests on un-reinforced soil samples. The horizontal displacement

required to mobilize the shearing stresses are measured along with the vertical

displacements of loading plates. Typical shear boxes used to determine the soil-

geosynthetic frictional properties are shown in Figures (2.8a-b). Test results are

usually expressed as the efficiency factor which is the ratio between the soil-

reinforcement interface friction angle and the soil friction angle (tan 6 / tan 0 ).

Different shear boxes have been utilized to evaluate the shear strength

at the soil-reinforcement interface. The sizes of shear boxes ranged from the

conventional ones (Richards and Scott, 1985) and small boxes of 4 in. long by

4 in. wide (Martin et al., 1984) to relatively larger shear boxes of 10 in. by 10

in. (Saxena and Budiman, 1985) and 12 in. by 12 in. (Degoutte and Mathieu,

1986; and Williams and Houlihan, 1987). The shortcoming of using small shear

boxes is due to their inability in measuring the large deformations required to

mobilize the interface friction between soil and geosynthetics. Moreover, the

large dimensions of most of the geogrid patterns do not allow using

representative samples in such small boxes. Myles (1982) and Miyamori et al.

(1986) used direct shear devices with the lower boxes of bigger dimensions

than the upper ones. The advantage of using a lower box of bigger dimensions


17

is to subject the reinforcement to a higher shear strain level without loss of

shear area. A comparison of dimensions of the shear boxes, reinforcements and

soil types used, and testing procedures is shown in Table (2.1).

Direct shear test results provide the local shear stress-strain relationship

at the soil-inclusion interface. However; in order to determine the pull-out

resistance of the reinforcement, pull-out tests have to be performed. Since pull-

out resistance integrates the non-uniform variation of the shear stress-strain

along the reinforcement, pull-out tests incorporate the reinforcement

extensibility effect on the interface parameters. Consequently, the interface

frictional parameters obtained from shear and pull-out tests can vary. Both tests

are associated with different testing procedures, loading paths and boundary

conditions. In selecting the most appropriate type of test it is essential to

adequately simulate field conditions. For soil reinforcement applications, pull-out

tests should be used since it is more representative of field conditions and

failure mechanism of embankments and slopes (Juran et al., 1988).


Normal Displacement

Normal Load

Load-Upper In terfacedUpper Soil

Fabric

• low er Soillo a d (Fabric)

H o riio n ta l d iio la c tm n U

(a). After Rowe et al. (1985).

to data logging systemi—reaction

structure \ [~

nrrrollers

rubber bag'' v ;

load cellpressurecells -reaction

structure to the shear load

—rollers

jachscalesoil sample

dmxlm xlm) 0.2 m

wm TTT

(b). After Palmeira and Milligan (1989).

Figure 2.8 Soil-Geosynthetic Direct Shear Test Devices.


19

References Device Description Reinforcement Type C )

Soil Type Notes

[1] Williams and Houlihan, (1987)

Shear box 12x 12 in’ .

- Trevira 1155 12)- Typer 3401 (2)- Nicolon 900 [1]• PVC membrane• Polyethelen [4]

- Ottawa 2 0 /3 0 • Concrete sand- Glacial till- Saprolite- Gulf coast clay.

- [if]-1 0 -4 5 deg.

- Rate - 0.01 in/min.

[2] Miyamori et el. (1986)

Lower box 47 .8 x 40 .8 cm1, upper

box 31 .6 x 31 .6 cm5.

- Poystsr needle- punched [1]

Poorly graded dry and wet sand.

- [ d / * i - .72-.7S dry sand,- .64- .86 wet sand.

- Rate - .5 mm/mi n

- [a) - 50 -500 Kpa

- Fabric glued to wooden block in lower box.

[3] Degoutte and Mathieu, (1986)

Shear box 30 x 30 x 3 0 cm3.

- Geotextile 11,2)- PVC 14]

• Sand [? l - 33° - Sandy clay

I ? ] - 33.5°

- [d j - 35-39 deg.

- Soil in one side of box

- I d ) - 200 -1200 Kpa

[4] Saxena and Budiman, (1985)

Shear box 1 0 x 1 0 in5.

• Celanese 600 [1] - Monsanto [2)

- Sandy clay, 45% Ottawa

- Saturated Lime stone balast.

- [ d | - 23 .8 - 27 deg. for C-34 - 17.2- 22 deg. for 600X

- Rate 0 .03 in/min

- [ol - 10-30 psi

[5] Formazin and Batereau, (1985)

Shear box - Non-woven- Woven

- Saturated sand - [ d i - 22 -35 dag.

[6] Richards and Scott, (1985)

Shear box 2x2 in5. -BidiumU-14 [21 •Terrafix [1) -Mirafi T 700 [1]

- Quartz angular sand [ p i - 38°.

- ( d ) - 31-41 deg.

(1) Reinforcement type:[1] wooven geotextile,[21 non-wooven geotextile,[31 geogrid,[4] other materials.

(2) $ « soil friction angle6 - interface friction angle.

Table 2.1 Comparison of Geosynthetic Direct Shear Tests, [from available experimental data]


20

• Table 2.1 [Continued] •

References Device Description ReinforcementType

Soil Type Notes

[7] Martin et al. (1984)

Shear box 10 x 10x 2 .54 cm3

- EPOM FML [4]- PVC FML [4]• Typer 3401 [2]- Polyfilter X [1]- Mirafi 5 00 X (1)

Concrete sand (?) « 24 deg.

-[<?)- 20 -30 deg.

•Soil in one side.

-Rate - .127 mm/min.

- [ o ] - 14-100 Kpa

[8] Myles, (1982) Upper box .1 m1, lower box .35 x .4 m!

- Needle-punched Polypropelene [2]

- Woven Polyster

Leighton Buzzard sand .4 -.85 mm

-[<*]- 36 -44 deg.

-Fabric glued to wooden plate in lower box.

-Rate «* 10-75 mm/min.

- |a ]« 30, 60 and 300 Kpa.

[9] Jacobsen, (1985)

Shear box 17.5 x 10 cm1, circular end plate to acheive shear at different angles.

- Fibertes needle- punched [2]

Quartz angular sand

-la] m 3-5 m of sand

(1) Reinforcement type:[1] wooven geotextile,[2] non-wooven geotextile,[3] geogrid,[4] other materials.

(2) <p “ soil friction angle6 « interface friction angle.

Table 2.1 [Continued].


21

In the pull-out tests, the rear end of the specimen is free while the front

end is clamped to the pull-out loading machine. Pull-out tests were used to

provide the load-displacement relationship at the facing of the geosynthetic

specimen and its pull-out resistance. Since no standard design for pull-out

testing devices exist, box dimensions and testing procedures differ for every

box. The dimensions of the box are usually chosen to reduce the boundary

effects. Figures (2.9 to 2.12) show typical pull-out testing equipment. A

summarized review of most of the available equipment and testing parameters

is presented in Table (2.2).

The review of the existing pull-out tests shows a large variety of testing

equipment and procedures which makes it difficult to compare test results. It

also indicates some significant limitations which the investigators tried to

overcome with different degrees of success. Most of these limitations are

related to:

(1) Most of pull-out tests were performed with controlled displacement

rates. Various pull-out rates were recorded in the literature (varying from 0.1

mm/min till 20 mm/min). Although Myles (1982) has studied the frictional

resistance of the geotextile interface in direct shear tests under different strain

rates (10-75 mm/min) and showed little sensitivity of test results, the effect of

displacement rate on pull-out results has to be evaluated. Moreover, very few

pull-out tests under load controlled mode are available in the literature.


22

*~ 1

Normal Displacement

Normal LoadUpptr Shear Box

K ■ ^wxxvvxxxxxvvsxxxxxxvxxvsxvwvwvll %

Figure 2.9 Pull-out Testing Device. (After Rowe et al., 1985)

C HornonFal Displacement

Outer Box

TOP VIEW

» p e c im e n w in d o w lood cell

CTS. jack

mo» orfm eoturem ent lo o t

Q lotil fo r th preeeure celle

S ID E V IE W

lid T £

.lo p r»»»ur» source

tube tor m ooneUc p ic ku p

_ .S P lc jm t!*— ----------i --------------d a m p load cell

---------n— p i— rvr-----

1 3 3Y O

Figure 2.10 Pull-out Testing Device. (After Holtz, 1977).


23

•9'

Figure 2.11 Geosynthetic Pull-out Box.(After Christopher, 1976)

POTENTIOMETER

T12 cm

SLEEVE5 5 c m

SUPPORT COUNTERWEIGHT

GEOTEXTILEREINFORCEMENT

Figure 2.12 Displacement Monitoring Along the Reinforcement inPull-out Test.

(after Juran and Chen, 1989)


24

References Pull-out Box [length x width x height] in3.

Reinforcement type (1)

Soil type Notes

[1] Elias, (1979) 3 6 x 36 x 18 Smooth and ribbed metal strips [4]

- Residual silt- Kaolin* clay- Wyoming sand

• Strain rate 0 .3 in/min.

- rigid front face.

[2] Tzong and Cheng-Kuang, (1987)

48 x 48 [sample width = 18 in.]

-Trivera 1127 [2] - Ottawa sand, density= 1 0 7 pcf

- Load controlled test.

- (<ri= 4 .3 4 psi.

[3] Brand & Duffy (1987)

1 2 x 1 2 [sample size 1 0 x 1 0 in’ .]

- Tensar SS2 [3]- Tensar SR2 [3]- Tensar SR3 [3]- Tensar TT2 [3]

• Bentonite clay - Strain rate 0.01 cm/min.

- [a] = 1 2 -4 8 Kpa

[4] Christopher, (1976)

48 x 30 x 18 - Tensar SS2 [3]- Signode TNX [3]

- Fontainbleau dense sand

- Strain rate 0.1 mm/min.

- Rigid front face with sleeves.

- [cr] = 440 , 660, 8 00 psf.

[5] Koerner, (1986)

18 x 18 x 8 - Geomembranes [4]-C 2 -6 0 0 [2]- Tensar SR2

- Well graded concrete sand

- Rigid front face.

[6] Ingold, (1983) 19.7x 11.2x 11.8 - Nelton 1168 [3]- Tenser SR2 [3]• Welded wire

mesh [4]- mild steel sheets

- Borham wood pit sand

- water filled air bag.

- rigid front face.

[7] Holtz, (1977) 4 2 x 10 x 11 - Woven polyster [1]

- Tullinge sand- G-12 rounded

sand.

(1) Reinforcement type:[1] woven geotextile[2] non-woven geotextile[3] geogrid[4] other material.

Table 2.2 Comparison of Geosynthetic Pull-out Tests, [from available experimental data].


25

Table 2 .2 [Continued].

References Pull-out Box [length x width x height] in3.

Reinforcement type (1)

Soil type Notes

[8] Rowe et al. (1985)

Pull-out and shear box on samples 6

x 6.

- Permealiner [1]- Mirafi [1]- Geolon 1250 [1]- Terrafix [2]- Tender SR2 [3]

- Loose silty sand. - Strain rate 1.0 mm/min.

- Rigid front face.

[9] Johnston, (1985)

54 x 5 4 x 36 - Tensar SR2 [3] - Medium to coarse sand.

- Strain rate 4 .4 mm/min.

[10] Anderson, (1984)

72 x 60 x 48 - Welded wire mesh [4]

- Silty sand- Washed sand- Pea gravel

- Strain rate 0 .8 mm/min

[11] Jewell, (1980)

Modified shear and pull-out box 10 x 6 x 6

- welded wire steel [4]

- Leighton Buzzard sand

- Strain rate 0.1 mm/min.

- Rigid front wall.

(1) Reinforcement type:[1] woven geotextile[2] non-woven geotextile[3] geogrid[4] other material.

Table 2.2 [Continued].


26

Tzong and Cheng-Kuang (1987) examined the pull-out interaction mechanism

under imposed loading-controlled conditions. Load-controlled pull-out tests are

necessary if the in-soil time-dependent interface parameters need to be

evaluated.

(2) The interaction between the soil and the box side walls can affect the

pull-out test results. The applied confining stresses can be partially carried out

by the side wall friction causing a reduction in the normal pressure applied at

the reinforcement level (Johnston, 1985). Anderson and Nielsen (1984)

minimized the interaction with the side walls by keeping the edge of the

specimen at 1.5 ft from the side walls. Jewell (1980) lubricated the walls with

silicon and covered them with a thin rubber membrane to provide frictionless

boundary. Several investigators confined the soil within a flexible membrane to

insure uniform distribution of normal stress by means of air pressure

(Christopher, 1976) or de-aired pressurized water (Ingold, 1983).

(3) The interaction between the reinforcement-soil system and the rigid

front wall can also influence test results. As the reinforcement is pulled out

from the box, the lateral earth pressure developed at the front face can result

in an increase in pull-out resistance. Christopher (1976) incorporated sleeves

around the pull-out slot to transfer the pull-out application point far behind the

rigid front wall. Other investigators ((Williams and Houlihan, 1987) used flexible

front face to minimize its effect. Johnston (1985) used pull-out box with the

front face removed to prevent interaction with the soil-reinforcement system.


27

(4) The thickness of soil differs in each box according to its clear height

{Table 2.2). If soil thickness is small, the interaction between the soil-

reinforcement system and the rigid plates at the top and bottom of the box may

restrain soil dilatancy and the mobilized shear resistance at the interface and,

consequently, affecting pull-out resistance. Brand and Duffy (1987) determined

pull-out resistance of geogrid in different thickness of clay. Their results show

that as the soil thickness increased, pull-out resistance decreased until a

minimum force state is reached.

(5) An apparent increase in the pull-out resistance of the inclusion was

recorded with the increase in soil relative density and normal pressure (Ingold,

1983; and Koerner, 1986). Johnston (1985) attributed recorded variations in

the normal pressure on the reinforcement level to the uneven sample

compaction. Different sample preparation and compaction procedures were

utilized to insure uniform soil density. Soil compaction was achieved by means

of electric jack hammer (Johnston, 1985), standard proctor hammer (Saxena

and Budiman, 1985), hand tamping devices (Elias, 1979) and by mechanical

tamping (Anderson, and Nielsen, 1984). A hopper with flexible tube was also

used to insure uniform soil placement in the box (Jewell, 1980).

(6) Several investigators (Christopher, 1976; Koerner, 1986; and Brand

and Duffy, 1987) clamped the reinforcement outside the box. The disadvantage

of this technique is that the unconfined front portion of the reinforcement

results in a variation of the effective interface area during the test.


28

(7) The existing pull-out testing equipment are usually instrumented to

monitor displacement, pull-out rate, and mobilized pull-out resistance at the

facing. Most of these equipment lack the proper instrumentation to monitor the

internal deformations along the interface. For extensible reinforcement,

monitoring the displacements at different locations along the inclusion is a key

element in a proper development of a load transfer mechanism. Jewell (1980)

used an x-ray device and placed short lead markers in the soil to detect the

internal deformations in the sand. However, the maximum penetration depth

(15 cm) of the x-ray limited the sample width used in this experiment.

Christopher (1976) used extensometers to measure the displacements at

various locations along the inclusion. Juran and Chen (1989) described a

mechanism where they hooked inextensible ,tell-tail, wires to the reinforcement

through slots at the rear wall (Figure 2.12) in order to monitor the

displacements at different locations along the inclusion.


29

2.3) CONCLUSIONS:

The soil-reinforcement interface properties can be measured in both

direct shear and pull-out tests. The interface friction coefficient is usually

presented as the efficiency factor which is the normalized value with the soil

to soil friction coefficient (tan SI tan 0). Efficiency factors ranging from 0.6 to

1.0 for geotextiles and values larger than one for geogrids were commonly

reported. The review of the available test results demonstrated that there exists

significant variations in the results obtained from both direct shear and pull-out

tests. Jewell et al. (1984) and Johnston (1985) found out that the frictional

resistances for various inclusions in dense sand were greater in pull-out tests

than those obtained from direct shear tests. Similar results were recorded by

Schlosser and Guilloux (1979). They attributed the higher values in pull-out

tests to the higher sample dilation which increases the local vertical pressure

on the reinforcement. Ingold and Templeman (1979) found equal values from

both tests for geogrids tested under low confining pressures. However, under

higher normal stresses, higher shear strength values were produced in pull-out

tests. Koerner (1986) reported higher shear resistance of geogrids in pull-out

tests. However, unlike the results of Ingold, he found out that under higher

confining pressures direct shear tests give higher shear resistance. Rowe et al.

(1985) found out that both tests give approximately equal values of shear

resistance for geotextiles tested in loose fine grained soil. Since loose sand has

low tendency for dilation, restrained expansion has no effect on normal


30

stresses. Figure (2.13) shows a comparison between the efficiency factors

obtained by various investigators from both pull-out and direct shear tests.

The different testing methods in both direct shear and pull-out boxes, the

associated boundary conditions and the corresponding soil dilatancy behavior

lead to different design parameters. Moreover, a fundamental difference

between both tests is that direct shear tests provide uniform shear stress-

displacement parameters; while these values are distributed non-uniformly along

the extensible reinforcement in pull-out tests.

The review of the available testing equipment has also showed that

many factors influence the measured interaction properties. These factors are

related to testing equipment and procedure, boundary condition, soil and

reinforcement type, soil placement and compaction, soil dilatancy behavior, and

overburden pressure. Several techniques are used in determining the effect of

these factors on test results; however, only with limited success.


31

t exf i iorctwr . :

o ♦□ ■

D R

C ollloa a t a l# (1980)

ItDMTSI-2(loOgltodiAal

d irection)

Pull-Out Direct Shear

NORMAL STRESS (PSD

Figure 2.13 Comparison of the Efficiency Factor Obtained from DirectShear and Pull-out Tests.

(After Juran et al. 1988)


CHAPTER 3

EQUIPMENT DESIGN AND INSTRUMENTATION

3.1) DESIGN CONSIDERATIONS:

Several pull-out facilities have been designed and used to determine the

soil-reinforcement interface properties of geosynthetics. A detailed review of

the available testing equipment is presented in the previous chapter to provide

guide lines for the design of the pull-out and direct shear testing equipment.

The review showed a large variety in the typed of available equipment and

testing procedures. It also showed some significant limitations which should be

considered in the design of the proposed equipment. These limitations are

mainly related to:

(i) Loading Scheme:

In most of the equipment reviewed, tests have been done under constant

displacement rates. Tests under various displacement rates should be

performed in order to evaluate their effect on the results. Moreover, the

equipment should provide the capability of conducting tests under load

controlled mode in order to assess the confined creep parameters of the soil-

reinforcement system. This testing scheme is of importance specifically for

geosynthetic materials which exhibit time-dependant deformations.

32


33

(ii) Rigid Wall Boundary Effects:

The rigid walls of the boxes influence test results by imposing boundary

conditions on the interface. The proposed equipment should provide, through

modular design, the flexibility in sample and box dimensions for the evaluation

of the boundary condition effects.

The side wall friction can partially carry out the applied normal pressure

and, thereby, affecting the confinement at the interface. Side wall friction can

be minimized by applying the normal pressure within a flexible membrane. Box

dimensions have also to be of adequate size to keep the confined specimen far

from box walls. In order to investigate the effect of the passive earth resistance

at the front wall, sleeves have to be incorporated into the design to insure that

the soil-reinforcement interaction is carried out far from the front face. Soil

pressure at the facing must be measured and a parametric study has to be

performed to determine the sleeve length at which this effect is minimal. In

order to evaluate the effect of top and bottom plates on the interaction

mechanism, tests with different soil thickness above and under the

reinforcement have to be performed.

(iii) Soil Placement and Compaction:

The frictional resistance of the inclusion is influenced by the relative

density of the soil. When interface shear stresses are mobilized during testing,

dense soil tends to dilate. As this dilation is restrained in the box, normal stress


34

increases at the vicinity of the inclusion. Furthermore, in geogrid reinforcement,

soil compaction increases the lateral resistance on the transversal elements

and, consequently, their shear resistance. A soil placement and compaction

procedure should be developed in order to facilitate sample preparation and

insure that the soil density is compacted uniformly to the desired density

throughout the box.

(iv) The Specimen Clamping Mechanism:

In pull-out tests, the clamping of the reinforcement to the loading device

outside the box leads to an unconfinement of the front portion of the

reinforcement. The unconfined elongation of the front part implies variation of

the interface area of the reinforcement during the test. The clamping

mechanism has to insure in-soil clamping of the specimen to maintain uniform

confinement and a constant interface area of the specimen during pull-out.

(v) Instrumentation:

The results of pull-out tests are influenced by the extensibility of the

inclusion. In order to evaluate the frictional resistance along extensible

materials, an interpretation method that incorporates the inclusion extensibility

has to be adopted. Accordingly, the instrumentation must be capable of

measuring the relative displacements at different locations along the confined

reinforcement.


35

3.2) EQUIPMENT DESCRIPTION:

The testing facility consists mainly of the following:

i) A pull-out box of dimensions 60 in. long, 36 in. wide and 36 in.

high.

ii) A large direct shear box with the lower part of dimensions 60 in.

long, 36 in. wide and 24 in. high, while the upper box is of dimensions

27 in. by 27 in. and 15 in. deep.

iii) A hydraulic loading system for each box, which is capable of performing

under both constant pull-out displacement-rate and constant pull-out load.

iv) A sand handling system; which consists of an elevated sand hopper and

a sand vacuum machine, that facilitates sand placement, removal and

compaction control for both boxes.

v) Instrumentation and data acquisition system to control and monitor

the input testing and response parameters (e.g. displacement rate, pull-

out load and displacements at different locations along the

reinforcement).

Figure (3.1) shows a side view of the pull-out box, the loading frame and the

sand handling equipment. Figure (3.2) shows a front view of the box, the

loading frame and the clamping plates. Figures (3.3) and (3.4) give the

longitudinal and cross sectional details of the pull-out box. The direct shear box

is shown in Figure (3.5), and its cross sectional details are shown in Figure

(3.6). Both boxes were constructed with ASTM A36 mild steel. The details of


36the cross sections of the testing equipment are shown in Appendix A.

U The Pull-out and Shear box Details:

The main elements of the pull-out and direct shear boxes are:

1) % in. bottom and side wall steel plates which can be assembled to

modify the length of the box. The boxes were assembled in bolted modular

units to adjust the length to 30 in., 45 in. and 60 in. and to adopt any increase

in the box dimensions in the future. For the tests performed in this study, box

length was kept at 60 in. For the pull-out box, the width of the box was chosen

to keep the standard sample, of 1 ft wide, at a distance 1 ft from each side of

the box wall to reduce the effect of side wall friction on the soil-specimen

interaction. The level of the hydraulic ram can be adjusted to allow pulling out

the inclusion at different heights in soil in order to evaluate the effect of soil

thickness on test results.

2) A modular front wall is designed of 4 in. by 4 in. rectangular beams. The

front wall contains a slot of 2 in. height to permit pulling the clamping plates

out of the box. In the pull-out box, the modular units permit an evaluation of

the effect of the rigid front boundary by using slots with variable opening sizes

and different facing types.

3) A modular rear wall of 2 in. by 4 in. rectangular beams. In the pull-out

box, five slots are located in the rear wall to permit the instrumentation for

displacement measurements along the reinforcement.


37

4) A loading frame , shown in Figure 13.1), which is bolted to the front face

of the box to support the hydraulic loading device.

5) An air bag of 2 in. thickness. It is used to apply the vertical overburden

pressure when inflated through the air pressure source. The pressure system

is able to apply a normal pressure up to 30 psi. A cover plate of % in. thickness

is used to confine the air bag above the soil. The cover plate is placed under

rectangular beams which are bolted to the box wall to enclose the air bag.

6) Sleeve plates, shown in Figure (3.3), of 4 inch width x Yi inch thick. The

plates are placed on the top and bottom of the slot in the front wall of the pull-

out box. The sleeves are placed at the facing to transfer the pull-out load far

behind the box rigid front face. The sleeves minimize the lateral stress transfer

to the rigid facing during pull-out which would result in an apparent increase of

the pull-out resistance of the inclusion. The sleeves are designed in modular

units to evaluate the minimum sleeve length required to eliminate the effect of

the rigid front face.

7) Two clamping plates, shown in Figure (3.2), of 1/8 in. thick. In the shear

box, the clamping plates are bolted to the upper box; while, in the pull-out box,

they are bolted to the reinforcement inside the box.


1) Pull-out Box2) Loading Frame3) Elevated Sand Container.

Figure 3.1 Side View of the Pull-out Box.


1) Loading Frame2) Clamping Plates3) Load Cell4) Geogrid Reinforcement.

Figure 3 .2 Front View of the Puli-out Box and Loading Frame.


40Beam 4*x4*

a

Air bag ,' cover plate!

01

Sleeve 13 Reinforcement Compactedi sand

Beam ,2*x4■01 Beam 4’x4

■

5 ’ - 0"11

Figure 3.3 Longitudinal Cross-Section of the Pull-out Box.

1 /2 ’ S t i f f *

n r r , . rrn XJ) J

Beam 4'x4"

V TT p\

• ‘ A ir b a g - / .

>

o o o o o

* rne Back wbII

.' * . *• • %

Beam 4*x4*

■+41+ --------------- 3 1 ^ ---------------------- X- 4 ^

a 5 r

itn

Figure 3 .4 Cross-Section of the Pull-out Box.


Figure 3.5 View of the Large Direct Shear Box.


42

Air bag

Geogrid

Compacted sand

5 ‘ - O'

Figure 3.6 Cross-Section of the Direct Shear Box.


43

ii) The hydraulic Loading System:

The pull-out load is applied by a hydraulic system. The system includes

three basic units:

i) hydraulic ram model Miller H67B (Figure 3.7). The ram is mounted on

the loading frame and applies the pull-out load on the clamping plates. The

hydraulic piston of the ram is 5 in. diameter and is able to apply 18 inch

maximum pull-out displacement. A servo valve is mounted on the ram which

controls the piston pull-out displacement rate.

ii) hydraulic power supply unit which consists of a hydraulic pump model

Miller of 5 HP and 20 gallon fluid reservoir capacity and a cooling system. The

details of the power unit are shown in Figure (3.8). The pump is able to operate

under two loading schemes, namely: (a) Pressure-control scheme; where the

pump fluid is controlled by a low pressure proportional control valve. The

pressure is measured directly in the hydraulic ram by a test gauge; and (b)

Velocity-control scheme; Where the pump operates under a constant pull-out

velocity with variable pressures up to 3000 psi. The constant fluid rate is

controlled mainly by the servo valve mounted on the hydraulic ram. The pump

is controlled to operate under either of theses schemes from the control box

unit.

iii) control box unit (Figure 3.8); where the commands to the hydraulic

pump can be sent manually by the control keys in the board or through the

computer when the control board is connected to the data acquisition system.


44

iii) Sand Handling Facility:

In order to minimize operator's effort and control the sand placement

procedure, a sand handling facility has been specially designed and constructed.

This facility is shown in Figure (3.9) and it consists of:

1) A heavy duty movable sand vacuum model 'Invincible': The sand is

vacuumed out from the box through a flexible hose and the dust is filtered out

to a dust container of 3 ft3 capacity. The vacuum machine stores the sand in

the elevated hopper.

2) Elevated sand hopper of capacity 54 ft3 of sand: The hopper is used

to store the sand when vacuumed from the box. The hopper supporting system

is designed to permit its movement above the two boxes. To place the soil back

into the box, the hopper is moved above the box and the sand is loaded back

by gravity through the flexible hopper outlet. The outlet is elevated to 55 in.

above ground in order to permit positioning the outlet above the box.


4 5

1) Hydraulic Ram2) Loading Frame3) Load Cell.

Figure 3.7 View of the Loading Frame.


A) Control BoxB) Fluid ReservoirC) Hydraulic RamD) Servo ValveE) Low Pressure Gauge.

Figure 3.8 Details of the Hydraulic loading System.


A) Elevated HopperB) Vacuum SystemC) Hopper Outlet.

Figure 3 .9 View of the Sand Handling Facility.


48

3.3) INSTRUMENTATION AND DATA ACQUISITION:

The instruments used to measure the interaction response and the input

testing parameters, for both of the pull-out and direct shear boxes, consist

mainly of:

1- load cell used to measure the pulling force applied from the hydraulic

loading system,

2- linear variable differential transformers (LVDT's) to measure the

displacement at the pull-out application point,

3- velocity transducers for measuring the front displacement rate of the

reinforcement,

4- pressure cells used to measure the internal earth pressure inside the

box,

5- manometer gauge to measure the normal pressure applied from the air

bag.

Moreover, in the pull-out box, LVDT's are placed at the back of the box to

measure the displacements at different points along the reinforcement. The

specifications of these instruments and the results of their calibration tests are

presented in Appendix A. The measured response is monitored by a data

acquisition system. The data acquisition system is used to:

i) translate the output response from the measuring instruments into digital

values in the computer,

ii) display and store the output data into the computer system,


49

iii) translate the operator's digital commands from the keyboard to analogous

values in the control box.

(i) Load and Pressure Measurements:

The pull-out force is measured by a load cell attached to the hydraulic

piston and the clamping plates as shown in Figure (3.7). The load cell of 20

Kips capacity operates when an excitation voltage is sent to the load cell from

the load cell conditioner. The conditioner, shown in Figure (3.10), also reads

and stabilizes the output signal and converts it to the data board in the

computer. The load cell response is a voltage output in millie-volts (mV) and it

corresponds to the magnitude of the applied load in pounds. Calibration tests

were performed on the load cell by applying different predetermined loads on

the load cell and monitoring its response. The results of these tests, which

show the relationship between the applied load in Lbs and the output response

in mV are shown in Appendix A.

When load-controlled pull-out tests are performed, the load can be either

manually controlled by the pressure valves in the hydraulic pump or computer

controlled through D/A channel of the data control board.

The earth pressure on the facing wall is measured using two earth

pressure cells model 'GeoKon 3650' with maximum pressure of 100 psi. The

pressure cells are 2 in. diameter and consist of two circular stainless steel

plates welded together and spread apart by a narrow cavity.


50

External pressure acting on the cell is balanced by an equal pressure

induced in the internal fluid by an excitation voltage. The excitation voltage of

10 volts DC is supplied from a stabilized power supply (Figure 3.10). The

pressure is converted by the pressure transducer into an electrical signal which

is transmitted to the data control board in the computer. Calibration tables to

convert this output response (in mV) to the magnitude of the applied pressure

(in psi) are provided by the manufacturer. These values were checked by

testing the cells in a triaxial pressure chamber. Predetermined incremental air

pressure was applied and the voltage output was monitored. The calibration

values of the earth pressure cells are shown in Appendix A. The earth pressure

cells are fixed on the facing wall in the locations shown in Figure (3.11).

(ii) Displacement Measurements:

The geogrid front displacement in the pull-out box and the upper box

displacement in the direct shear box are monitored by means of Linear Variable

Differential Transformers (LVDT's). These LVDT's are mounted on the loading

frame and are shown in Figure (3.7). The LVDT's used are of 18 in. length and

stroke length of ± 10 in. An excitation DC voltage of 15 volts is sent by means

of stabilized power supply. Displacement is monitored as the core rod moves

inside the LVDT causing an output voltage response equivalent to the

displacement. The relationship between the core rod displacement and the

output voltage is provided by the manufacturer.


51

These calibration tables are presented in Appendix A and they are utilized

in the A/D channels of the data control board to convert the voltage values to

digital output.

For the geogrid extensible reinforcement, non-uniform shear stresses and

displacements are developed along the reinforcement with the large shear being

mobilized at the front portion of the inclusion (Christopher, 1976; and Juran

and Chen, 1988). Therefore, It is necessary to monitor the displacements at

different locations along the reinforcement in order to determine the

reinforcement length which is effectively mobilized under a specific level of pull-

out load. In order to measure the displacements at different locations in the

confined inclusion, five LVDT's are mounted at the level of the inclusion on a

rear table in the pull-out box. The displacement rods of the LVDT's are

connected to the geogrid by 'tell-tale' inextensible wires through slots in the

box rear wall. The wires are connected to the transversal ribs of the geogrid

and kept stretched by means of counter weights through the rear table. A detail

of the setup at the rear table is shown in Figure (3.12) and a view of the

displacement instrumentation of the geogrid at the rear table is shown in Figure

(3.13).

(iii) Displacement-rate Measurements:

A velocity transducer is mounted on the loading frame of both boxes to

measure the front displacement-rate. The velocity transducer measures the rate


52

of the core rod displacement when it is moved inside the transducer shaft. An

excitation voltage of 5 volts is sent by the DC power supply. The output

voltage is calibrated by applying several predetermined velocities and the

equivalent output voltages are monitored. The displacement-rate is measured

for the following purposes:

(i) to compare the displacement-rate at the front during pull-out with the

command value sent by the computer when the test is conducted under

a displacement-rate controlled mode.

(ii) to measure the variation of front displacement-rate when the test is

conducted under constant loading scheme.

When displacement-rate controlled tests are performed, a command is sent to

operate the hydraulic ram under a constant velocity. The command is sent to

the hydraulic pump through the D/A channel of the data control board in the

computer. The velocity control is calibrated by operating the pump with

different command values (in volts) and measuring the output voltage response

of the transducer. The velocity transducer output is then equivalent to the

displacement-rate calculated from the time-displacement measurement. In

displacement-rate controlled tests, the pressure of the hydraulic pump is kept

constant at a high value (1000 psi) in order not to affect the velocity control

procedure.


53

(iv) Data Acquisition:

The output response parameters (i.e. displacement, velocity, load and

pressure) measured, respectively, by the LVDT's, velocity transducers, load

cells and pressure cells; are transmitted to display and storage in the computer

through a data acquisition system. A schematic diagram of the instrumentation

and the data acquisition system is shown in Figure (3.14) and a view of the

data acquisition control system is shown in Figure (3.15).

A data translation board (model DT-2801 AD/DA) is plugged into one of

the computer system extension slots. The data translation board includes 8

differential (16 single-ended) A/D channels and two D/A channels. The dual

function of the board is to:

(i) translate the digital commands, sent by the operator, to analogous ones

through the D/A channels,

(ii) translate the analogous output data of the tests to digital data to display

and storage through the A/D channels.

The data input and output are simultaneously monitored and stored in the

computer. The data translation board is connected to the measuring

instruments through a screw terminal board model DT707. The screw terminal

board is shown in Figure (3.10). The connection scheme of the screw terminal

board is:

(i) each of the measuring instruments is connected to one of the A/D

channels of the screw terminal,


54

(ii) the screw terminal A/D channels are connected to the control box in

the hydraulic system. These channels translate the commands to the

hydraulic loading system for either hydraulic ram velocity control or

pressure control.

A schematic diagram of the screw board showing the assigned channels to

each instrument is shown in Figure (3.16). The data acquisition board is

programmed and controlled from the computer keyboard through the software

"Labtech Notebook'. The software supports the data control management and

organizes data display and storage.


A) Stabilized Power SupplyB) Screw Terminal BoardC) Load Cell Conditioner.

Figure 3.10 Data Acquisition Supporting Instruments.


cover Plate

air bagFront MallJ— c e l l #1

j - r - c e l l #2clamping plates

sleeve

eogrid-

Figure 3.11 Locations of the Earth Pressure Cells.

r

y-

/ /

. /

Box Rear Hall

core rod

aR ear T a b le

A

©LVDT

unextenslble Mire

counter weight

Figure 3.12 Detail of the 'Tell-tale' and the Rear Table.


^ocfuced

uCJQ.

CD

CD

in

h-o

Pull-Out Box

Figure 3 .14 Schematic View of the Data Acquisition System.


Figure 3.15 View of the data Control System.


60

• (g)

LVDT#3LVDT«4

LVDT #5 .LVDT#2

Velocity

Load CelTLVDT at Facing

LoadVelocity Control

Control'

© s / n E

Figure 3.16 Connection Scheme of the Screw Terminal Board.


CHAPTER 4

TESTING PROCEDURE AND DATA INTERPRETATION

4.1) INTRODUCTION:

Several pull-out and direct shear tests were conducted in order to

evaluate the performance of the testing facility, the effect of testing parameters

and to provide the data base for the development of an analytical data

interpretation procedure. These tests are:

(1) Unconfined extension tests: These tests were conducted on the geogrid

specimens under constant extension-rates. Results were compared with the

index test results provided by the manufacturers. These tests are aimed to

evaluate the equipment performance, the accuracy of the control and

monitoring system, and the material behavior in the unconfined state.

(2) Pull-out tests: Constant displacement-rate pull-out tests were conducted

under the same testing conditions (e.g. confining pressure, soil density,

boundary conditions, and pull-out rate) to evaluate the equipment reproducibility

and the pull-out performance of the geogrid. Tests were also conducted under

variable testing conditions to evaluate the effect of testing parameters on the

pull-out resistance.

(3) Direct shear tests: These tests were conducted to evaluate the shear

stress-strain characteristics at the soil-geogrid interface. Results of these tests

were also used to evaluate the interface parameters calculated from the pull-out

61


62

tests using the proposed data analysis procedure.

(4) Load-controlled pull-out tests: Stepped load-controlled pull-out tests

were performed to determine the unconfined and confined time-dependant

parameters of the geogrid. In these tests, loads were applied incrementally and

maintained constant during specific times. A data analysis procedure was

adopted to predict the critical creep load and strain. Long term (500 hour) creep

tests were performed to evaluate the confined creep strain of the geogrid under

the critical creep load.

The testing methodology and interpretation of test results are discussed

in this chapter. The procedures used in specimen preparation, soil compaction,

and calibration of soil density are presented.

4.2) SOIL COMPACTION AND SPECIMEN PREPARATION:

Soil compaction and relative density significantly affect pull-out test

results (Knochenmus, 1987; and Juran et al., 1988). In dense sand, as soil

particles are displaced in the vicinity of the reinforcement, the soil tends to

dilate. This dilation is restrained by the surrounding soil. This results in a normal

stress concentration at the soil-reinforcement interface which increases the pull-

out resistance of the soil.

For a given relative density, soil placement and compaction can have a

significant effect on the results obtained from the pull-out box. Uneven soil


63placement and compaction would result in a sample with higher tendency for

arching or non-uniform soil dilation. Previous studies (Rowe et al., 1985; and

Johnston, 1985) showed that non-uniform sample dilation leads to uneven

normal pressures along the inclusion. Inadequate soil placement and poor

compaction of soil layers would result in differential settlement of the inclusion,

which may lead to an incorrect measurement of the interface pull-out

resistance. The compaction process should be adequately adopted for each

type of soil in order to simulate, as closely as possible, the compaction process

used in the field. The compaction techniques and the calibration procedure used

to determine the amount of compaction effort needed to produce the required

relative soil density are discussed herein.

i) Soil Placement:

The sand used in this study is a uniform blasting sand. Its grain size

distribution is shown in Figure (4.1). The shear stress characteristics under

different confining pressures, determined from the conventional direct shear

test, are shown in Figure (4.2). The maximum and minimum soil density tests

were conducted according to ASTM D4253-83 and D4254-83, respectively.

The values of 99 pcf (1.58 t/m3) and 110.9 pcf (1.77 t/m3) were obtained,

respectively, for minimum and maximum soil density.

The sand is placed into the pull-out box by pouring it from the elevated

hopper through a flexible outlet. The elevated hopper is moved during pouring


64to permit uniform filling of sand along the box. The sand is placed in four layers

of 6 inch each, leveled and compacted to the desired relative density. Figure

(4.3) shows the sand placement scheme in the box. Compaction was applied

manually using a vibrating electric hammer. A predetermined number of blows

was applied on each layer in a regular pattern.

After compaction, the density was measured with a nuclear density

gauge model 'TROXLER 3440'. Figure (4.4) shows the soil density

measurement in the pull-out box. The gauge rod was embedded 8 in. to 12 in.

into the soil and the density is measured at 9 different locations to insure

uniform density distribution at each layer.

ii) Compaction Control:

In order to simulate the compaction process commonly used in the field,

a compaction process which consists of coupling vibration and hammering

effect has been adopted. A vibrating electric hammer was modified to hold a

compacting plate of 5 inch diameter allowing manual compaction of the soil

layers.

Calibration tests were performed in order to establish the compaction

effort and sequence required to achieve the required relative density. For the

purpose of performance evaluation study, calibration tests were performed with

two different sand placement procedures, namely:

(a) Variable falling height: In this procedure, the level of the outlet pipe in


65the hopper was fixed. This leads to different sand falling height from the hopper

outlet to each of the sand layers in the box.

(b) Constant falling height: In this procedure, the hopper outlet is connected

to a long flexible hose to keep sand falling height constant from the outlet to

each layer.

Under each of these two soil placement methods, different compaction efforts

were applied to the sand layers; namely,

1- No compaction: the hopper outlet was moved above the box to permit

uniform fill and the soil layers were leveled with minimum disturbance to

achieve minimum soil density.

2- 40 blows per layer: Soil was compacted by applying 40 blows uniformly

at each layer.

3 - 11 0 blows per layer.

4- 200 blows per layer.

At each layer, density was measured after compaction and before placing the

following layer. Density was measured again when upper sand layers were

removed in order to assess the effect of the weight of the upper layers on the

density of the compacted sand. The details of compaction schemes and results

of density measurement are shown in Appendix B. The relationship between

the applied number of blows and the resulting soil density, for the two cases

of variable and constant sand falling heights, are shown in Figures (4.5-a and

b), respectively. These curves can be used to estimate the number of blows per


66

layer necessary to obtain a desired soil density. The figures show that, under

the same compaction effort, the first method of sand placement (i.e. variable

falling height) induces a looser state of sand at the lower layers. The constant

falling height procedure (which maintains the falling distance constant above

each layer) allows better control of compaction and results in soil densities

independent of falling height. The effort induced in leveling each layer should

also be the same in order to reduce the effect of sand placement on the

density.

The results in Figure (4.5) indicate that after compacting a specific layer,

the successive placement and compaction of the following layers results in an

additional increase in sand density which should be taken into consideration

when the number of blows is to be determined. A systematic and consistent

sand placement, leveling and compaction techniques should be used in order

to achieve a reproducible initial state of soil density.

iii) Reinforcement Preparation:

Two different types of geogrids, commonly used in soil reinforcement,

were tested. These geogrids are: (a) geogrid 'TENSAR SR 2\ and (b) Geogrid

'CONWED'. For both types of geogrids, standard specimens of 1 ft width x 3

ft length were tested in the pull-out box. In the large shear box, the specimens

had the same size as the upper box (i.e. 27 in. width x 27 in. length).


67In the pull-out box, the standard specimen width allows keeping its edge

at 1 ft from the side walls of the box to reduce boundary effects. The

specimens were bolted to the clamping plates and placed over 1 ft of sand. A

sleeve of 1 ft length was used around the clamping plates at the front wall. The

sleeve transfers the pull-out application point far behind the front wall in order

to reduce the effect of front rigid boundary on test results. Figure (4.6) shows

the placement of the geogrid in the pull-out box.

The displacement distribution along the confined geogrid is measured by

five LVDT's placed at the 'tell-tail' table. Figures (4.7-a) and (4.7-b) show the

locations of the displacement measurements along the geogrids 'Tensar' and

'Conwed', respectively. The LVDT's are connected to the transversal ribs of the

geogrid through non-extensible wires. The wires extend through 1/2 inch

plastic tubes embedded in the sand to prevent the frictional resistance of the

sand to the movement of the wires. A sand layer of 1 ft thick is placed above

the reinforcement and the confining pressure is applied from the air bag.


m

ah*•*■* ait

ill!f i l la:cs

Figure 4.1 Grain Size Distribution of the Blasting Sand.


Shea

r St

ress

(K

N/m

)

80

60

40

20

Dry density- 106 pcf (1.7 t /« 3)

Or - 60%

A * zA Arm 65 KN/ *A A A

AA

A A

„ . . cr - 10 KH/mZ

i_________ i_________ i_______0 1 2 3 4 5

H o rizo n ta l D is p la c e m e n t (m m )

Figure 4 .2 Direct Shear Test Results on the Sand.


70

HDPPER

-L a ye r 4 6*

Sleeve L a ye r 3 \

6 '

Laye r 2 6 '

L a ye r 1 . ♦ *•* ,6"

f

5 '-0 '

Figure 4 .3 Sand Placement in the Pull-out Box.


Figure 4 .4 Soil Density Measurements in the Pull-out Box.


Soil

Dens

ity

(pcf

) So

il De

nsity

(p

cf)

Layer 4

10B

Layer 3 (down)

Layer 3

!(up)

Layer 2

102

100

9 8 200150

Bl100

Num ber of Blows5 0

ows

(a) Variable Falling Height Procedure,

110

108

106

102Layer 4

Layer 3 (down)

Layer 3 (up)

Layer 2

100

9 85 0 100

Num ber150

Num ber of Blows200

(b) Constant Falling Height Procedure,

Figure 4 .5 Calibration of Sand Density with Compaction.


duced with

74

Locations of Displ. Instrumentation

Geogrid TENSAR SR2

Clamping Plates^3 ’ - O’

(a) Geogrid TENSAR SR2'.

9 ribs ,

i — A -

12 ribs

3

3. ft

c lam p ing

plate

(b) Geogrid "CONWED'.

Figure 4.6 Instrumentation of Geogrid Specimens in Pull-out Box.


754.3) DISPLACEMENT-RATE CONTROLLED TESTS:

i) Unconfined Extension Tests:

Unconfined extension tests were performed on the geogrids in order to

evaluate the material behavior in the unconfined state and the equipment

performance and its reproducibility in the displacement-rate controlled mode.

The unconfined stress-strain test results were compared with the index tests

provided by the manufacturers.

Unconfined extension tests were performed on geogrids 'Tensar SR2'

and 'Conwed'. In these tests, the front face of the specimens were bolted to

the clamping plates while the rear ends were clamped to rear wall of the pull-

out box. A computer command was sent to operate the hydraulic pump with

a constant pull-out velocity of 2 mm/min. The front displacement, the velocity

and the load were monitored during the test. Figures (4.8-a) and (4.8-b) show

the unconfined stress-strain relationships of geogrids 'Tensar' and 'Conwed',

respectively. The index test results of both geogrids, provided by the

manufacturer, are plotted in their respective figures. The figures show that the

measured unconfined response compares fairly well with the index results of

the geogrids.

ii) Pull-out Tests:

Pull-out tests were conducted on both 'Tensar' and 'Conwed' geogrids

in order to evaluate their pull-out resistance and interface parameters. In these


76tests, standard testing conditions were maintained constant in order to evaluate

the reproducibility of pull-out test results. A confining pressure of 7 psi (48.2

KN/m2), an average soil density of 104 pcf (1.67 t/m3), a soil thickness of 2 ft,

a displacement-rate of 6 mm/min, and a sleeve length of 12 in. were kept

constant in the pull-out standard tests. In these tests, geogrid specimens of 1

ft width and 3 ft length were tested. Figures (4.6-a and b) show the dimensions

and the points of displacement measurements along the geogrids tested in the

pull-out box.

Figure (4.9) shows the load-displacement relationships from a series of

pull-out tests on 'Tensar' geogrid samples tested under the standard testing

conditions. The figure shows that the results yield similar pull-out resistance,

interface stiffness modulus, and front displacement at peak. Figure (4.10)

shows the results of a series of tests conducted on 'Tensar' geogrid under a

confining pressure of 7 psi, average soil density of 105.5 pcf (1.69 t/m 3), and

a displacement rate of 10 mm/min. The figure shows that an increase in the

soil density and displacement-rate results in an increase in the peak pull-out

resistance of the geogrid. The effect of variation of testing parameters (e.g. soil

density, confining pressure, displacement-rate, and boundary conditions) on the

pull-out response of the geogrids will be discussed in detail in chapter 5.

Figure (4.11) shows the pull-out response from a series of tests on

geogrid 'Conwed' of standard specimen sizes and under confining pressures of

7 psi, and average soil densities of 106 pcf (1.7 t/m 3).


77

The measured front displacements at different nodes along 'Tensar'

geogrids are shown in Figure (4.12). The figure shows that the slopes of the

displacement curves at the front correspond to the controlled input velocity.

The displacements at different nodes along the geogrid are monitored by means

of the LVDT's at the rear table. Figures (4.13) and (4.14) show, respectively,

typical time-nodal displacement relationships for 'Tensar' and 'Conwed'

geogrids. The results show the progressive load transfer mechanism along the

geogrid. The interpretation of the results in these figures shows that:

(i) At the front node (i.e. at node 0), the slope of the front displacement

verses time is constant and is equal to the controlled pull-out displacement rate.

(ii) At early stages of pull-out response, the difference in the slopes of

nodal-displacement curves indicates progressive movement of the geogrid

nodes during testing. This progressive load transfer leads to higher extension

at the front part of the geogrid with practically no strain at the rear nodes.

(iii) The slopes of time-nodal displacement curves become practically equal

at the later stages of testing, indicating that the geogrid extension is fully

mobilized along its length. At this stage, the geogrid moves as a rigid body

without any further extension at this stage. The specific point on the

displacement curve where the slopes become practically equal indicates the

occurrence of the slippage failure of the reinforcement which is attained at the

peak pull-out resistance.

To illustrate the progressive load transfer mechanism along the geogrid


in pull-out tests, the nodal displacements are normalized with respect to the

front displacement (i.e. node 0). The normalized displacements at different

loading levels are plotted along the geogrid length in Figures (4.15) and (4.16)

for geogrids 'Tensar' and 'Conwed', respectively. The figures show that at an

early stage of testing (at loading level 40% of peak pull-out load), the front

nodes of the geogrid (nodes 1 and 2) develop their displacement response while

no displacements are practically monitored at the rear nodes inodes 4 and 5).

At this early stage of pull-out loading, the grid experiences an extension

response at the front part without slippage and the strained length of the

geogrid is its effective adherence length. At a later stage of pull-out testing, the

geogrid reaches its peak strength, the rear node undergoes displacement, and

the geogrid experiences a combined extension-displacement response. This

combined response is mobilized along the reinforcement till, at a later stage of

post peak response, the normalized displacement curve becomes practically

linear. The linear displacement along the geogrid specimen demonstrates that

the geogrid experiences a rigid body movement without further extension.


Load

(K

N/m

) Lo

ad

(KN

/m)

79

Unconfined Extension

Geogrid ’Tensar SR2*

Index Test

□ Test *1

• Test * 2

a Test #3

10 15 2 0

Axial Strain (X)2 5 3 0

(a) Geogrid 'Tensar'

5 0


Geogrid ’Conwed’4 0

3 0 o •□ •

20Index Test

A Test: C0N-U110 - a • Test: C0N-U2

O Test; C0N-U3

0105 150

Axial Strain (%)

(b) Geogrid 'Conwed'.

Figure 4 .8 Unconfined Extension Test Results on Geogrid.


80

60

Confining pressure= 7 psi (48.2 Kn/m2) Soil density= 104 pcf (1.67 i /m 5)

z * • . “ o*^ 4 0 - p . - . r * .□ m■oao

0 ci 1 o Test: A—3

20 P a Test: A -5

• Test: A—7

a Test: A -8

50 100 150 200

Front Displacement (m m )

Figure 4.9 Pull-out Test Results on Geogrid 'Tensar SR2'.

80

Confining pressure= 7 psi (48.2 KN/m2)

Soil density= 105.5 pcf (1.69 t /m 3)60 -

E

. • • • •° a •

□ • ° •a A ° Q ,* AOA °• A * Q

TJOO*m

0 •— 0

A°

8 00

□ Test: C -2

▲ Test: C—3

• Test: C -4

50 - 100 150

Front Displacement (m m )200

Figure 4.10 Pull-out Test Results on Geogrid 'Tensar SR2'.


80

60

E\z

"DDO

20

00 10 2 0 3 0 4 0 5 0 60

Front Displacement (m m )

Figure 4.11 Pull-out Test Results on 'Conwed' Geogrid.

Confining pressure= 7 psi (48.2 KN/m 2)

Soil Dens!ty= 106 pcf (1.7 t /m 3)

A | 0

A tf>

1 ■ Test: Con-1

O Test: Con-2

A Test: Con-3

I______ I______ l______ l______ l____

A A ^ A A

at O A

160

140

120

EE 100

CoE 80

! 60 a

40

20

■ □ Test—A5

▲ A Test—A7

• o Test—A8

Displacement at Peak g

o9 A

Face Displ. g 0 D

. s ° °8 □8 aa □ #an

on " a6 0 # A "6 ° • A ■

- * A ■

aO

A ■• A m

• Am • Am

• Am • Am

1 ; ■ Node-3 Displ.

a * t * | Soip-M — 1--------

i t *

200 400 600 800 1,000 1,200 1,400Time Ised

Figure 4.12 Measurement of Front Displacement in Pull-out Tests.


82

200

150 Front Displ.

Node (1) Node (2)

aVBf)0 ti1 5

100

0 500 1000 1500 2000T im e ( te c )

Figure 4.13 Time-Nodal Displacement Relationship for Geogrid 'Tensar SR2' in Pull-out tests.

100.12 ribs9 ribs,

clomping

plote8 0

o.

Node# 2

TJ 20

0 500 1,000 1,500 2,000 2,500Time (sec)

Figure 4.14 Time-Nodal Displacement Relationship for 'Conwed' Geogrid in Pull-out Tests.


83

Test A—5 Test A—7o A □© 0.8

At Displ.= 100 mm"5.0.6

*o2 0.4

At Displ.= 50 m'm

At Displ.= 20 mo 0.2

Geogrid Nodes

Figure 4 .15 Normalized Displacements Along the 'Tensar' GeogridSpecimen.

1 +Geogrid Conwed

Test: Con—1© 0.8

a 0.6 At Displ 50 mm

TJ2 0.4

At dfsftl 30 mm (Peak)-

O 0.2

At displ. 20 mm

Geogrfd Nodes

Figure 4 .16 Normalized Displacements Along the 'Conwed' GeogridSpecimen.


84

4.4) DIRECT SHEAR TESTS:

In the direct shear test, the clamping plates are bolted to the upper shear

box which is permitted to move between the lower box and the confining

beams on its top (see Figure 3.5). Calibration tests were conducted without the

sand in the box and the resistance induced by the friction between the box

walls and the upper beams was measured.

Direct shear tests were performed on sand to sand interface, without the

geogrid, in order to determine the shear stress-strain characteristics of the

sand, the soil internal friction angle (0) was equal to 29.8 degrees under

confining pressure of 7 psi and average soil density of 106 pcf. Direct shear

tests were then performed with the geogrid sample at the interface. Geogrid

samples of 72 in. by 72 in. were placed on the bottom of the upper shear box.

An interface friction angle (£) equals 32 degrees was obtained under the same

testing conditions. Figure (4.17) shows the effect of geogrid reinforcement on

the frictional interaction at the interface. An efficiency factor (tan 01 tan 6) of

1.1 is obtained under the specified confining pressure and soil density. This

value is similar to results obtained by other different investigators (Koerner,

1986; and Ingold, 1983) where efficiency factors more than one were recorded

at the interface of 'Tensar SR2' geogrid.


1

tntnto£ 0.8 CO

D£ 0.6 OZ\tnS 0.4k.

cn

g 0.2-Ctn

o0 10 20 ' 30 40 50

Displacement (m m )

Figure 4.17 Effect of Reinforcement on the Frictional Resistance inDirect Shear Test.

Direct Shear Test

Geogrid ’Tensar SR2’

Soil Density= 106 pcf (1 .7 t /m S)

# — . .y nBBaaQ,n,; , * . 1

. * D°°o0m& ° d d " D d d b□

D

□ sand /sand Interface

] ■ sand/geogrid interface


86

4.5) LOAD-CONTROLLED TESTS:

When geosynthetic reinforced samples are subjected to sustained loading

conditions, the strain response consists of an elastic strain followed by a time-

dependant creep strain (Figure 4.18). The main design concerns for the long

term stability of the reinforced earth structures are:

(i) to predict the long term creep displacement under a constant pull-out

load,

ii) to evaluate the critical creep load (or creep strain) below which creep

rupture is unlikely to occur.

In order to determine the long term creep displacement under a constant

load, load controlled pull-out tests are performed and displacement is monitored

with time.

The critical creep pull-out load can be determined by subjecting the

specimen to a stepwise increasing load over a time interval. This interval can

be determined at each load level when the deformation is stabilized. The critical

creep load can be established following a procedure similar to that used for

ground anchors (Christopher et al., 1989) which is illustrated in Figure (4.19).

In this procedure, the measured front displacement, for each load, is plotted

versus Log time (T). An upward concavity of the creep curve indicates an

accelerated creep failure. The slope of the displacement versus Log (T) is

plotted against the applied pull-out load to determine the critical creep load Tc.


87

The creep strain induced by a specific loading can be determined as the

sum of the strains produced by the previous load increments (Vyalov, 1986).

This concept of strain superposition is illustrated in Figure (4.20). Stepped load

tests has the advantage that one specimen is tested under the different loading

levels which reduces the effect of testing parameters on the results.

In the load-controlled pull-out tests, the loads are incrementally applied

to the inclusion and maintained constant during a specified period. The induced

load, displacement and velocity at the front are recorded during the test. Three

types of load-controlled tests were performed; namely:

(i) load-controlled extension tests on the unconfined geogrid,

(ii) Stepped load-controlled pull-out tests on the confined geogrid to

determine the critical creep load,

(iii) long term pull-out tests under a constant load to evaluate the creep pull-

out displacement of the confined geogrid.

i) Unconfined Load-Controlled Extension Tests:

Unconfined extension tests were performed on geogrid 'Tensar' of 1 ft

width and 3 ft length. The purpose of these tests was to evaluate the time-

dependent unconfined behavior of the geogrid under stepped loading conditions

and to evaluate the performance of the testing facility under this testing mode.

A stepwise increasing load was applied to the samples starting from 10% of

the maximum strength of the unconfined geogrid. The load was increased to


88

40% and 70% of the maximum load. Figure (4.21) shows the unconfined time-

dependent strains in these tests.

ii) Stepped Load-Controlled Pull-out Tests;

Stepped load-controlled pull-out tests were conducted on the "Conwed'

geogrid. In these tests, the specimens were subjected to different loading levels

defined as percentages of the maximum pull-out resistance (Tmax) obtained

from displacement-rate controlled tests. Figures (4.22) and (4.23) show the

front displacement verses time relationship of the geogrid under confining

pressures of 7 psi and soil density of 107 pcf. In these tests, the specimens

were subjected to stepwise loads increasing from 20% Tmax to 85% Tmax at

time intervals of about 1800 min. for each step.

In order to interpret the critical creep pull-out loads, the measured front

displacements, for each loading level, were plotted versus log (time) as shown

in Figure (4.24). In these curves, the displacement-rate curves indicate

stabilized creep with time till a loading level of about 70% Tmax where creep

pull-out rate starts to accelerate. The maximum displacement rates at each load

level were plotted against the normalized pull-out load (T/Tmax) in Figure (4.25)

in order to estimate the critical creep load (Tc). The long term creep

displacement under the critical creep load (Tc) was evaluated in a long term

(500 hr) pull-out test under the critical creep load.


89

iii) Long Term Pull-out Tests:

Long term pull-out tests were conducted on 'Tensar' geogrid specimens

subjected to pull-out load equals (Tc) in a soil dry density of 107 pcf. The value

of Tc was estimated from Figure (4.27) and was taken as 65% of the

maximum pull-out resistance of the geogrid. Figure (4.26) show the creep

displacement of the geogrid specimen under confining pressures of 7 psi. The

Log (displacement-rate) versus log (time) relationship of the geogrid is plotted

in Figure (4.27). The Figure shows an upward of the creep curve indicating an

accelerated creep.


RupturePrimaryCreep

SecondaryCreep

TertiaryCreep

Time

Figure 4.18 Illustration of Creep Concept.

ton n

0 . 6 5 ^

0.4 Tr

Log T 0.5 - 1.m

Figure 4.19 Determination of the Critical Creep Load. (After Christopher et al. 1989).


91

(o)

(b)r

<c)

r

T T “ wr

Figure 4.20 Concept of Strain Superposition in Stepped Loading Tests.(After Vyalov, 1986).


92

20Test: U -1

Test: U—2

at 70% Tmax

(?)

at 40% Tmax

at 10% Tmax

400- 200 Time (m in)

300100

Figure 4.21 Unconfined Load-Controlled Tests on Geogrid 'Tensar'.


93

20

TEST: C0N-C1

15

851 Tux

10

CL

BOX Tb b x5

00 50 0 1.000 1,500 2.000 2 ,5 0 0Time (m in)

Figure 4 .22 Stepped Load-Control Test on Geogrid 'Conwed'.

20

Test: C0N-C2

15

B5X Tmax

10

u70S Tmax

CL

5

S5X Tmax

00 500 1,000 1 ,500 2.000 2 ,5 0 0Tim e (m in )

Figure 4 .23 Stepped Load-Control Test on Geogrid 'Conwed'.


Confined Creep Test

0.05 ik

0.030.02

.6055 Tmax

D 0.00550J5 Tmax

Q 0.0032055 Tmax

0.002100 200 300 500 1,000Time (min)

2,000

Figure 4 .24 Log(Dipl. rate) versus Log (Time) in Creep Tests on 'Conwed'

0.03c.E\ 0.025EE 0.020)oL_

0.015cwE0)u 0.01oa.no 0.005c

Geogrid Conwed

0.6 Tmax / •

m /

f /- •

» /........•#••••.............

1 * i_. i0.2 0.4 0.6

T/Tm ax0.8

Figure 4.25 Determination of the Critical Creep Load.


EJ,mievsoa

JBQ.os

20

TEST: CREEP-1

15

10

5

010001001 10

■nine (hour)

Figure 4.26 Results of Creep Test on Geogrid 'Conwed'.

10'

Test: CREEP-1

[ at 65X Tmax ]

CD COD oa>

■SooaDaootoocMBafea

aGeogrld: Conwed

10" *

10* 10*101TIm (hour)

Figure 4.27 Results of Creep Test on Geogrid 'Conwed'.


CHAPTER 5

EFFECT OF TESTING PARAMETERS ON THE INTERACTION MECHANISM

5.1) INTRODUCTION:

The soil-reinforcement interaction involves mainly three basic load-

transfer mechanisms:

(i) lateral friction on the soil-reinforcement interface (as in geotextiles,

strips and bars),

(ii) passive earth pressure on the transversal elements of geogrids, welded

wire meshes and bar mats,

(iii) particle interlocking as in the geotextiles and geogrids.

Load transfer in most of the geosynthetic materials is a combination of these

mechanisms. Mobilization of these mechanisms requires different magnitudes

of displacements that would substantially affect the interaction performance.

The relative contribution of each mechanism to the total pull-out resistance

depend on many factors, namely:

(i) material properties (e.g. its type, geometry, extensibility and creep

properties),

(ii) soil characteristics (e.g. its relative density, shear strength and grain

size distribution),

(iii) loading conditions (e.g. overburden pressure, displacement-rate and

testing boundary conditions).

96


97

The major objective of the performance evaluation study is to assess the

sensitivity of test results to the changes in the testing parameters and establish

a data base for the development of a reliable interpretation procedure. For this

purpose, a parametric study was conducted and the effects of the main testing

parameters (i.e. geogrid type, displacement rate, confining pressure, rigid box

boundaries, soil compaction and relative density) on the soil-geogrid interaction

response were investigated. The tests performed for this parametric study are

shown in Table (5.1). The table shows the main testing parameters for each

test. The detailed results of these tests are presented in Appendix C.


Test Material Sample width [cm]

Confiningpressure[KN/m2]

Displ. rate

[mm/min]Soil

density[t/m3]

Soilthick[cm]

Sleevelength[cm]

Notes

(a) Displacement-rate Extension Tests:#1 Tensar-SR2 30 Unconfined 3 — — — sample#2 9 9 30 extension 3 — — — clamped#3 9 9 30 tests. 3 — — — at end.

Con-Ul Conwed X3022 30 Unconfined 2 — — —9 9Con-U2 9 9 30 extension 2 — — —

Con-U3 9 9 30 tests. 2 — — —Con-U4 9 9 30 2 — — —Con-U5 9 9 30 2 — — —Con-U6 9 9 30 2 — — —

(b) Displacement-rate Pull-out Tests:Cl Tensar-SR2 30 48 10 N/M(*) 60 30 standarC2 9 9 30 48 10 1.69 60 30 tests.C3 9 9 30 48 10 1.69 60 30C4 9 9 30 48 10 N/M 60 30C5 9 9 30 48 10 N/M 60 30A3 9 9 30 48 6 1.66 60 30A4 9 9 30 48 6 1.67 60 30A5 9 9 30 48 6 1.67 60 30A7 9 9 30 48 6 1.67 60 30A8 9 9 30 48 6 1.67 60 30

(*) Not measured

Table 5.1 List of Pull-out Tests.

99

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Table

5.1

[Continued]

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Table

5.1

[continued]

101

5.2) EFFECT OF TESTING PARAMETERS ON INTERACTION RESPONSE:

(i) Reinforcement Type:

The mobilized shear resistance at the soil-reinforcement interface

reinforced soil systems depends on the type of reinforcement. In case of

geogrids, the interface shear strength is primarily mobilized, as illustrated in

Figure (5.1), by the skin friction and passive resistance against transversal ribs.

For coarse grained soils, the openings in the geogrids may allow soil particles

to interlock between the ribs and, thereby, increasing its shear strength. The

skin friction between the soil and the geogrid depends on the type of soil and

the surface roughness of the geogrid. The contribution of soil resistance at the

transversal ribs to the overall shear strength of the geogrid depends on many

factors such as confining pressure, geogrid geometry and grain size distribution

of soil. The effect of soil friction at the soil-geogrid interface is shown in Figure

(5.2). The figure shows the results of pull-out tests on 'Conwed' geogrids

performed with and without their transversal ribs. These tests were performed

under confining pressures of 7 psi (48.2 KN/m2) and 14 psi (96.4 KN/m2). The

figure shows that an average value of 75% of the overall pull-out resistance of

this specific geogrid is attributed to the frictional resistance on the longitudinal

ribs.


Load

(K

N/m

)

a) Friction Between Soil £ Surface

Figure 5.1 Load Transfer Mechanism for Geogrids.

60

60

4 0

20

0

Figure 5.2 Pull-out Tests on Geogrids With and Without the Transversal Ribs.

Geogrid ’Conwed’

Soil Denslty= 106 pcf (1.7 t/m ^ )(96 kh/ b1)

7 psi 148.2 KN/» Ii— • —• --------- * »

0 n (96 KN/sr)

7 psi (48.2 KN/m*-)

o O No—Transversal Ribs

■ • With Transversal Ribs

10 ,-20 30Front Displacement (m m )

40


103(ii) Displacement Rate:

Currently, no standard testing procedure that specifies the displacement-

rate for pull-out tests exists. For unconfined extension tests, ASTM Designation

D4595-86 recommends a standard rate of strain of 10 ± 3 % per min for wide

strip method test. Rowe and Ho (1986) performed wide strip tests on geogrids

at strain rates from 0.2% to 10% per min. They showed that the geogrid

tension strength varies according to the applied rate of strain. It should be

noticed that the wide strip test is an index test under unconfined conditions

and does not represent the reinforcement response under confinement. Under

confining conditions, reinforcement experiences its peak pull-out strength at

much lower strain levels and test duration should be relevant to the observed

fabric behavior in field conditions.

For pull-out tests, most of the previous studies were conducted under

different displacement-rates. Table (5.2) shows the values of pull-out

displacement rates used by different investigators. The displacement rates

varied from 0.1 to 15 mm/min. In order to investigate the effect of

displacement-rate on pull-out response, tests were performed on geogrid

'Tensar' under four different pull-out displacement-rates, while maintaining the

other testing parameters constant. These pull-out displacement rates are:

i) 0 .033 mm/sec (2 mm/min)

ii) 0 .10 mm/sec (6 mm/min),

iii) 0 .18 mm/sec (10 mm/min),


104

iv) 0.33 mm/sec (20 mm/min).

In these tests, sample dimensions of 1 ft (.3 m) width and 3 ft (1 m) long, an

applied normal stress of 7 psi (48.2 KN/m2), and an average soil density of 104

pcf (1.67 t/m3) were maintained. Figure (5.3) shows the pull-out response of

the geogrid under these different displacement-rates. The Figure shows that the

increase in the displacement-rate from 2 to 20 mm/min results in a reduction

in the peak pull-out resistance and interface stiffness modulus.

Figure (5.4) shows the effect of displacement-rate on nodal

displacements. The plotted displacements were measured at the front point

(node 0), and at 2 ft (0.61 m) from the geogrid front end (node 4). The slopes

of the curves at node (0) are constant and consistent with the input commands

of displacement-rates. The slopes of the displacement curves at node (4) show

that higher displacement-rates result in a faster mobilization of shear stresses

at the rear nodes.

In order to evaluate the effect of the applied displacement-rate on the

displacement distribution along the geogrid, displacements along the nodal

points are normalized with respect to the front displacement (i.e at node 0) and

are plotted along the geogrid in Figure (5.5). The Figure illustrates that, at peak

pull-out level, higher displacement-rates result in a more uniform shear strain

mobilization along the inclusion.

The effect of displacement-rate on the geogrid pull-out resistance and its

interface stiffness modulus are shown, respectively, in Figures (5.6) and (5.7).


105The figures show that the peak pull-out resistance and the interface stiffness

modulus appear to be less sensitive to the changes of displacement-rates under

6 mm/min. These results suggest that in order to reduce the effect of

displacement-rate on pull-out results, pull-out displacement-rate should not

exceed 6 mm/min.


1 0 6

Box Testing Method Testing rate

- Caltrans Box (Johnston, 1985) [54"x36"x20"]

displacement rate controlled pull- out on Tensar SR2

2%/min.(4.4 mm/min)

- V. Elias Box (Elias, 1979) [36"x36"x18"]

displacement rate controlled on metal strips

0.1 in/min (2.5 mm/min)

- STS Box (Christopher,86) [48"x30"x18"]

displacement rate controlled pull- out on Tensar SR2.

0.1 mm/min

- Utah State U.Box (Anderson,1984)

displacement rate controlled on wire mesh.

0.8 mm/min

- Drexel U. (Koerner, 1986)

displacement rate controlled pull- out on tensar SR2.

0.25-15.2 mm/m

- U. of Grenoble (Gourc, 1980)

displacement rate controlled pull- out on Tensar SR2.

0.25 in/min (6 mm/min)

Table 5.2 Comparison Between Testing Strain Rates for Different Pull-Out Boxes,

(After Knochenmus, 1987)


N .

Zac■o«o

Geogrid: Tensar SR22

Confining pressure = 7 psi (48.2 KN/m ) Average soil density = 104 pcf (1 - 67 t/ms)

60. -

30.

OO-OOo 2 mm/min

6 mm/rrflPimm 10 mm/min

50 „ -100 150Front Displacement (m m )

200

Figure 5.3 Effect of Displacement-rate on Pull-out Response.

200

6 mm/min

Node §0 10 mm/min

20 mm/min

500 , - 1.000Time (sec)

Node #4

• 20 mm/min

A 10 mm/min

■ 6 mm/min 1.500

Figure 5.4 Effect of Displacement-rate on Nodal Displacements.


108

Geogrid: Tensar SR2

Displacement at Peak Loada> o.B

q _ 0.6

0.4

20 m m /m in

10 m m /m in

6 m m /m in

0.2z

0 1 2 4 53Geogrid Nodes

Figure 5.5 Effect of Displacement-rate on Displacement Distribution.


3Q.J£ID

30

15

Geogrid: Tensar SR22

Confining pressure = 7 psi I 48 KN/m )

Average soil density = 104 pcf ( 1.67 t/m3}

s 10 15Displacement-rate (m m /m in )

20

Figure 5.6 Effect of Displacement-rate on Pull-out Resistance.

n 2000

M3#■<3o 1500M10O*c

ui 100001uto

01

Geogrid: Tensar SR2

Confining pressure = 7 psi (4B KN/m2)

Average soil density = 104 pcf (1.67 t/m5 )

5 10 15

Displacement-rate (m m /m in )20

Figure 5.7 Effect of Displacement-rate on Interface Stiffness Modulus.


110(iii) Soil Compaction and Relative Density:

The frictional resistance of the inclusion is highly influenced by the soil

relative density (Juran et al. 1988). Dense soils tend to dilate when shear

stresses are mobilized along the reinforcement interface. As the dilatancy is

restricted by the surrounding soil, this results in a normal stress concentration

at the interface and consequently an apparent increase of the pull-out

resistance.

To evaluate the effect of the relative soil density on the pull-out

response, tests were performed with the soil compacted to different relative

densities under the same confining pressure; namely:

i) average soil density of 102 pcf (1.62 t/m3),

ii) average soil density of 104 pcf (1.67 t/m3),

iii) average soil density of 105.5 pcf (1.7 t/m3),

iv) average soil density of 108 pcf (1.73 t/m3).

Figures (5.8) and (5.9) show the effect of soil density on geogrid 'Tensar'

tested, respectively, with pull-out displacement-rates of 6 mm/min and 10

mm/min. In these tests, the geogrid was subjected to constant confining

pressure of 7 psi (48.2 KN/m2). The figures show that an increase in the soil

relative density results in an increase of the peak pull-out resistance and

interface stiffness modulus of the geogrid.

Soil compaction increases the lateral earth pressure on the geogrid

transversal elements and the mobilized frictional resistance at the interface.


111

Consequently, it restrains geogrid movement and slippage. Figure (5.9) shows

that when the soil was compacted to a relative density of 85% , the restrained

geogrid movement resulted in a significant increase in the interface stiffness

modulus and the pull-out strength. In this test, the reinforcement failed by

successive rupture of its longitudinal ribs at a very low front displacement.

The nodal-displacement distribution along the geogrid are normalized with

respect to the front displacement and plotted in Figure (5.10). The figure shows

the effect of soil density on the mobilized shear distribution along the geogrid.

The higher soil density results in a higher shear stress concentration at the

vicinity of the point of application and, ultimately, rupture failure at the front

element. Low soil density leads to a more uniform mobilization of the interface

shear stresses along the reinforcement.


112

80

60

Es2

40

TDOO

20

Geogrid 'Tensar SR2’

Confining pressure= 7 psi (48.2 KN/m 2)

Pull-out Velocity= 10 m m /m in

• Soil Density = 105.5 pcf (1.69 t /m 5)

□ Soil Density = 102 pcf (1.62 t /m * ). J50 100 150

Face D isp lacem ent [mm]200

Figure 5.8 Effect of Soil Density on Pull-out Response of 'Tensar' Geogrid.

120

90

Geogrid: Tensar SR2

Confining pressure = 48 KN/m2

Pull-out velocity = 6 mm/min

z

T3IDO

% Breakage Failure

60

30• Soil Density- 10B pcf (1.73 t /s r j

a so il Density- 104 pcf (1.67 t /« 5l

50 -100 150Front Displacement (m m )

200

Figure 5.9 Effect of Soil Density on Pull-out Response of 'Tensar' Geogrid.


113

■ A verage soil density = 1 . 6 7 t /m '

o Average soil density = 1 . 7 t / m s

▲ Average soil density = 1 . 7 3 t / m :v 0.8

"a 0.6

TJ • 0.*

O 0 .2

D i s p l a c e m e n t a t P e a k L o a d

Geogrid Nodes

Figure 5.10 Effect of Soil Density on Displacement Distribution Alongthe Geogrid.


114(iv) Confining Pressure:

The effect of confining pressure on the frictional resistance of the

reinforcement has been demonstrated by several investigators (McGown et al.,

1982; and Juran and Chen, 1989). For geogrid reinforcement, the confined

elongation during pull-out testing is restrained by the passive resistance of the

soil and particles interlocking within the transversal geogrid elements. The

restrained elongation results in an apparent increase of the tensile strength and

of the stiffness modulus.

Figure (5.11) shows the effect of the confining pressure on the pull-out

response of 'Tensar' geogrids measured under confining pressures of 5 psi (34

KN/m2) and 7 psi (48 KN/m2). These tests were performed with displacement-

rate of 6 mm/min and sand average density of 104 pcf (1.67 t/m 3). The results

show that an increase in the confining pressure results in an increase the pull-

out resistance of the geogrid. Although the increase in the confining pressure

reduces soil tendency to dilate, it results in an increase in the passive soil

resistance to the transversal ribs and, consequently, an increase in the geogrid

pull-out resistance.

The effect of soil confinement on the extension behavior of geogrid

'Tensar' is illustrated in Figure (5.12). In the unconfined tests, the samples

were clamped to the box rear wall; while, in the confined tests, the specimens

were tested under normal stresses of 7 psi (48.2 KN/m2) and 10 psi (70

KN/m2), and soil density of 108 pcf (1.74 t/m3). The figure shows that an


115increase in the confining pressure results in an increase in the pull-out

resistance and the geogrid ultimatly demonstrates successive ruptures of the

longitudinal ribs at lower strain levels.

Figure (5.13) shows test results of geogrid 'Conwed' when tested under

confining pressures of 7 psi (48.2 KN/m2), 14 psi (96.4 KN/m2) and 21 psi (140

KN/m2). The Figure shows similar effect of confinement on the pull-out

resistance of the geogrid. The results in Figures (5.12) and (5.13) show that

'Conwed' geogrid demonstrates rupture at a much higher confining pressure

than 'Tensar' geogrid. This is mainly due to the small thickness of 'Conwed'

transversal ribs (1 mm), compared to the relatively thick transversal ribs of

'Tensar' (4.5 mm), which requires higher confinement to mobilize passive soil

resistance at the transversal ribs.

These results suggest that the design criteria for geogrid reinforced soil

structures should take into account the in-soil confined extension properties

derived from pull-out tests rather than the material properties obtained from

unconfined extension tests. Although a lower (and consequently more

conservative) peak strength is obtained from unconfined tests, the unconfined

behavior of the geogrid is substantially more ductile and its strain level at peak

is higher than that obtained under confined conditions. This may results in non

conservative design values for the admissible geogrid extension and the related

structure displacements.

The normalized displacement distribution along the 'Tensar' and


116

"Conwed' geogrids under different confining pressure are shown, respectively,

in Figures (5.14) and (5.15). The figures illustrate the effect of the confining

pressure on the mobilized load transfer along the reinforcement.

The soil-geogrid interface shear stress is more uniformly mobilized along the

geogrid under low confining pressures. The increase in the confining pressure

restrains the geogrid displacement and results in a higher mobilization of the

soil-geogrid interface shear stresses near the pull-out application point, a lower

mobilization of the shear stresses at the rear, and consequently a shorter

effective adherence length.


11780

Geogrid ’Tensar SR2’

Soil Density = 104 pcf (1.67 t/r r? )

60

E

40 - °

-aao

Confining pressure 48 KN/rr

n o .' D O ,

■ ■20

Confining pressure 34 KN/m

50 ,100 150Front Displacement (m m )

200

Figure 5.11 Effect of Confining Pressures on Pull-out Resistance of 'Tensar'Geogrid.

150

120

Geogrid ’Tensar SR2’B r e a k a g e Soil D e n s ity = 1 .7 5 t / m 5

•Uoo

r/ i \u— \90 f •• • C on f. p re s s u re 7 0 K N / mI *.I ’• \I •I C o n f . p r e s s u r e 4 8 K N / m

U n c o n f in e d E x te n s io n

100 120Front Displacement (m m )

140

Figure 5.12 Effect of Confining Pressure on the Extension Behavior of 'Tensar'Geogrid.


Load

(K

N/m

)

118

90

75

60

45

30

15

00 10 , - 20 30 40

Front Displacement (mm)

Figure 5.13 Effect of Confining Pressure on the Pull-out Resistance of'Conwed' Geogrid.

.1 4 5 KN/m2

"■■■■ 96 KN/m2Ji n O O O O 00 °°° ® O O o O OH OO 0 °

- J P ° ..............- .........................................48 KN/m2- ■ u (>

■■ o■ o _ ,

o*„ °.F °»

* Geogrid 'CONWED*

S o il Density - 1 .7 t /m 3


119

i *

cO 0.8

euaO. 0.6wo

V 01N"oEO 0.2z

* * \ Confining pressure 5 psi

(34 KN/n)

Confining pressure 7 piff -... (48 KN/ma)

Displacement at Peak

0 1 2 3 4 5Geogrid Nodes

Figure 5.14 Effect of Confining Pressure on the Displacement DistributionAlong the 'Tensar' Geogrid.

Confining Pressure

2.

a 0.8

o 0.2

48 KN/m

96 KN/m‘

145 KN/m

Geogrid ‘Conwed’Displacmement at Peak

j --------------------- 1--------------------- L0 1 2 3 4 5

Geogrld Nodes

Figure 5.15 Effect of Confining Pressure on the Displacement DistributionAlong the 'Conwed' Geogrid.


1 2 0

(v) Sleeve Length:

The interaction between the soil-inclusion system and the rigid front face

of the pull-out box can affect the measured pull-out resistance. As the

reinforcement is pulled-out of the box, lateral pressure develops against the

rigid front face and results in an apparent increase of the geogrid pull-out

resistance. This rigid boundary effect can be reduced by means of a sleeve

incorporated around the slot in the front wall. These sleeves transfer the pull-

out application point inside the soil mass far beyond the rigid front

wall. The proposed standard ASTM testing procedure (1987) recommends the

use of sleeves at the entrance with a minimum length of 500 mm (20 inch) in

order to transfer the force into the soil far from the box front face.

In order to incorporate the effect of sleeve length in reducing the rigid

front face effect, tests were conducted with different sleeve lengths, namely:

i) no sleeve at the front face,

ii) 8 in. sleeve length,

iii) 12 in. sleeve length.

These tests were performed with soils of average density of 104 pcf (1.67

t/m3) and under a confining pressure of 7 psi (48.2 KN/m2). Figure (5.16)

shows the effect of the sleeve length on the 'Tensar' geogrid pull-out

resistance. The results demonstrate that the increase in the sleeve length

results in a reduction in the earth pressure developed at the front rigid wall and,

consequently, a reduction of the pull-out resistance. The lateral pressure


121developed at the facing was measured in these tests using two earth pressure

cells fixed on the box rigid front wall at a 3 and 8 in. above the sleeve (see

Figure 3.11). Figure (5.17) shows the development of lateral earth pressure on

the front wall when no sleeves are incorporated . The initial lateral pressures

recorded in both cells were approximately the same and were equal to 5 psi.

The lateral earth pressure in the upper cell (cell#2) substantially increased to 15

psi, while lateral pressure of 7 psi was recorded at the lower cell. Figures

(5.18) and (5.19) show the earth pressure recorded when sleeves of 8 and 12

in., respectively, were incorporated in the front face. The figures show a slight

increase of earth pressure at cell#2, while the pressure readings in cell#1

remained practically constant.

The results demonstrate the significant affect of the sleeve length on the

earth pressure measured at the front face. The use of the 8 in. sleeve results

in a substantial reduction of the earth pressure on the front wall. The increase

of the sleeve length from 8 to 12 in. has practically no effect on the lateral

earth pressure. The relationship between the sleeve length and the developed

lateral pressure on the front wall is shown in Figure (5.20). The figure shows

that the increase in the lateral earth pressure during pull-out becomes negligible

when a sleeve length of 12 in. is used. The effect of sleeve length on the

displacement distribution along the geogrid is shown in Figure (5.21).


122

120Geogrid: Tensar SR2

Confining pressure = 48 KN/m

No SLeeve60

TODO

Sleeve = 8 in.

30 r,•o

Sleeve = 12 in.

0 50 100 150 200Front Displacement (m m )

Figure 5.16 Effect of Sleeve Length on the Pull-out Response of 'Tensar'Geogrid.


123

140

1050i_3M*»0

D. 70 '

Geogrid: Tensor SR2

Confining pressure = 48 KN/m2 a a a

- Soil Density = 1.67 t/m3 * *A A Cell #1

A , . * «

,.' A* 4A

truuuu □ Q O DOOOOOD DODOOO

Cell #2

100 200 300Time (sec]

400 500

Figure 5.17 Development of Earth Pressure on the Front Wall in Case ofNo-Sleeve.

601ME

1 48

0k.

S 36 0

01=5 24 ?c£ »2

Li.

00 100 200 300 40 0 500

Time (sec)

Figure 5.18 Development of Earth Pressure on the Front Wall in Case of8 in. Sleeve.

Geogrid: Tensar SR2

Confining pressure = 48 KN/mz

Soil Density = 1.67 t/m3Cell #1

A A A A A A A

_________A A --------------- 1 — rA 4 . ^ ' V ' a A A A A A A

o o a o a*

»

A - A A

□ a

/ □o o o a

Cell #2

-----------------------------1----------------------------- 1_________ i___ i


Fron

t W

all

Pre

ssur

e (K

N/m

2)

12460

Geogrid: Tensar SR2

Confining pressure = 48 K N /m48

Cell #1■ I T ----

36

Cell #2D O D

DD

24

5000 100 200 300 400Time [sec]

Figure 5.19 Development of Earth Pressure on the Front Wall in Case of12 in. Sleeve.


125

140Geogrid: Tensar SR2

£Conf. pressure = 48 KN/m

Measurements at end of Tests.Ez:

a)i_

0) 70 a>i_

CL

Cell #1O

35-4->co Cell § 2

Li_

Sleeve'Length [in]

Figure 5.20 Effect of Sleeve Length on Lateral Earth Pressure.

1* -

Geogrid: Tensar SR2

Displacement at Peak Load® 0.8

a> 0 .4

Sleeve = 12 in.

Sleeve = 8 in. No Sleeve

Geogrid Nodes

Figure 5.21 Effect of Sleeve Length on Normalized Displacement Along theGeogrid.


1 2 6

(vi) Thickness of the Soil :

The interaction between the soil, the geogrid and the rigid boundaries

above and below the reinforcement can affect the pull-out test results. These

rigid boundaries can cause an increase in the normal stresses in the vicinity of

the geogrid surface if the soil thickness is small enough to restrain the soil

dilatancy. Moreover, friction can develop between the top and bottom rigid

boundaries and the soil which affects the soil-geogrid mobilized shear stress.

The proposed standard ASTM testing procedure (1987) recommends a

minimum thickness of soil layer of 100 mm (2.5 inch) above and below the

geosynthetic. In order to investigate the effect of the thickness of the soil on

the pull-out response and to evaluate the proposed ASTM recommendations,

pull-out tests were performed in soil samples of different thickness; namely:

i) soil thickness 10 cm above and 10 cm under the geogrid,

ii) soil thickness 10 cm above and 30 cm under the geogrid,

iii) soil thickness 30 cm above and 30 cm under the geogrid,

iv) soil thickness 40 cm above and 40 cm under the geogrid.

The above tests were performed on geogrid 'Tensar' under confining stress 7

psi (48.2 KN/m2) and average soil density of 104 pcf(1.67 t/m 3). The effect of

soil thickness on the pull-out response of the geogrid is shown in Figure (5.22)

and the effect of the soil thickness on the normalized displacement along the

geogrid is shown in Figure (5.23). The figures show that:

i) The decrease in soil thickness results in an apparent increase in the pull-


127out strength of the geogrid.

ii) The decrease in soil thickness leads to higher shear stresses mobilized at

the rear due to the developed frictional resistance.

iii) The increase in the soil thickness from 30 cm to 40 cm above and under

the geogrid does not have any significant effect on the pull-out resistance

of the geogrid.

These results lead to the conclusion that:

i) The proposed ASTM standard test method for measuring pull-out

resistance (1987), which recommends a total pull-out box height of 20

cm, would lead to high pull-out response values due to the effect of the

box boundaries above and under the reinforcement on the test results,

ii) The soil thickness of 30 cm (1 ft) above and under the reinforcement (a

total of 2 ft soil thickness) is enough to eliminate the effect of the upper

and the lower boundaries on the pull-out response.


Load

(K

N/m

)128

120

90

60

30

10 cm above Sc 10 cm under the geogrid




B P B B P -B -B -b -B -I

Geogrid: Tensar SR2

Soil Density = 1.67 t/m

Confining Pressure = 48 KN/m

50 100 150Front Displacem ent (m m )

200

Figure 5.22 Effect of Soil Thickness on Pull-out Response.

■ 10 cm above Sc 10 cm under the geogrid

A 30 cm above Sc 30 cm under the geogrid

O 40 cm above and 40 cm under the geogrid

Geogrid: Tensor SR2

Displacement at Peak Load. ____ I________1_____

1Geogrid Nodes

v5

Figure 5.23 Effect of Soil Thickness on the Normalized Displacement Alongthe Geogrid.


CHAPTER 6

MODELING LOAD TRANSFER MECHANISM.

6.1) INTRODUCTION:

The design of soil-reinforcement systems requires an evaluation of the

pull-out performance of the reinforcement with respect to three basic criteria:

i) The pull-out resistance of each reinforcement should be able to resist the

working tensile force in the reinforcement with an adequate factor of

safety.

ii) The front soil-reinforcement displacement, required to mobilize the pull-out

resistance, should be smaller than the allowable design displacement.

iii) The long term creep load, and creep strain, should be smaller than the

corresponding critical creep value.

In order to obtain the pull-out design parameters that satisfy the above criteria,

a load transfer model based on the interpretation of pull-out test results has to

be developed. This model should be able to evaluate the in-soil mechanical

properties of the reinforcement (i.e. the in-soil stress-strain behavior of the

reinforcement and its confined extension), and the soil-reinforcement properties

at the interface (i.e. the interface shear stress-strain relationship and the

reinforcement pull-out resistance).

The large number of factors that affect the interface properties raises

difficulties in developing an interpretation procedure from pull-out test results.

129


130

These factors are usually related to:

i) testing: such as apparatus and sample dimensions, boundary conditions,

strain rate, and the applied confining pressure,

ii) soil characteristics: such as its relative density, dilation properties, and

fine grain portion of the soil,

iii) reinforcement: its type, geometry and extensibility.

The major concern in the interpretation of the pull-out tests results pertains to

the effect of the reinforcement extensibility on the interaction pull-out

parameters. Inextensible reinforcement (such as metallic strips) moves as a rigid

member in the soil and the soil displacement required to mobilize the interface

lateral friction is very small (Schlosser and Elias, 1978; and Schlosser et al.,

1983). Such reinforcement develops a constant shear stress distribution along

its length. While; for extensible reinforcement, its non-uniform extensibility

results in a decreasing shear displacement distribution along its length. The

interface shear stress is therefore not uniformly mobilized along the total

reinforcement length. As the reinforcement extensibility depends upon its

length, pull-out interface parameters become a function of the specimen length

in pull-out tests. Therefore, the scale effect has to be evaluated in order to

extrapolate pull-out test results to the reinforcement length in the field.

Several investigators (Jewell et al., 1984; Stewart et al., 1977) defined

a soil-reinforcement resistance coefficient 6 to be used in the shear strength


131

relationship:

r = ov tan 6 ....................................... (6.1)

where, r = shear stress at the interface,

av = normal stress at the reinforcement level,

The pull-out resistance can then be calculated from the relationship:

Pr = 2 . y d tan 6 . L. (6.2)

where, Pr = pull-out resistance in force per unit width of the reinforcement,

Y = soil density,

d = thickness of soil above reinforcement,

L„ = length of reinforcement resisting pull-out force.

The disadvantage of this method is that it ignores the non-uniform shear stress

distribution along the extensible reinforcement. It assumes an even distribution

over the length and defines 6 as the slope of the pull-out stress versus normal

stress. Bonczkiewicz et al. (1988) modified the previous approach by

distributing the mobilized shear stress over the actual portion of the elongated

sample. They used wire extensometers to determine the sliding length of the

reinforcement as the pull-out progress. The pull-out shear per unit ares can then

be calculated as:

mobilized loadPull-out stress = ..................................................... (6.3)

actual strained section x 2

The FHWA manual for the design of reinforced soil structures (Christopher et

al., 1989) recommended a modified form of Equation (6.2), to include the


132

effect of extensibility, in determining the pull-out resistance:

Pr = F \ a . ov . L. . c (6.4)

where,

F’ is a pull-out resistance factor obtained from pull-out tests performed in

the specific backfill,

c is the reinforced unit perimeter [c = 2 for grids and sheets],

a is a scale effect correction factor.

The scale effect correction factor represents the nonlinearity of the Pr-L„

relationship due to extensibility. For inextensible reinforcement, a is

approximately 1. For extensible reinforcement, a can be determined from the

equation:

a = ....... (6.5)

where ravg and rp are, respectively, the average and ultimate interface lateral

shear stresses mobilized along the reinforcement. The disadvantage of this

empirical formula is that a depends primarily upon the interface shear strain

softening and not, explicitly, upon the extensibility of the reinforcement.

However, the manual recommends performing pull-out tests on reinforcements

with different lengths or using an analytical or numerical load transfer model to

obtain the scale effect factor.

Solomone et al. (1980) proposed another methodology for obtaining the

soil reinforcement resistance coefficient 6 by using the slope of the relationship


133

between the pull-out force and the mobilized length of the reinforcement (Kr).

The mobilized length of the reinforcement was found by assuming that

movement at a wire gauge at a specific location along the reinforcement

indicated pull-out at that point. The slope Kr can then be implemented, to

determine 6, in the relationship:

Kr / 2b = av . tan 6 (6.6)

where b is the reinforcement width. The limitation of this method is that it

assumes a uniform distribution of the pull-out force along the reinforcement, in

the form of the slope Kr, which is not the case in extensible reinforcement.

The review of the previous approaches shows the limitations of most of

these methods regarding the assumptions of uniform shear stress-strain

distributions along the reinforcement. The concept of a uniformly mobilized

interface shear stress which can be used for inextensible reinforcement can not

be adequately used for the interpretation of pull-out tests in extensible

reinforcement. In order to develop an appropriate load transfer mechanism, an

instrumentation scheme should be implemented in pull-out tests to measure the

interface shear strain at different points along the reinforcement.

Different investigators evaluated the interface shear strain distribution in

extensible reinforcement by obtaining measurements along the length of the

pull-out specimens. Figure (6.1) shows the displacement distribution along

instrumented geotextile specimen in a large box pull-out test (after Tzong and


134

Cheng-Kuang, 1987). Figure (6.2) shows the pull-out resistance distribution

along the length of the instrumented geotextile specimen in the pull-out box

(after Bonczkiewizc et al., 1988).

Leshchinsky and Field (1987) measured the load on both ends of non

woven geotextile specimens using the pull-out test apparatus shown in Figure

(2.3). The relationships between the applied load versus front and rear

displacements are shown in Figure (6.3). A second degree polynomial function

was assumed to interpolate the displacement U(x) at any location x along the

length of the fabric:

U(x) = K 0 + K 1 . x + K 2 . x 2 (6.7)

A linear distribution of the elongation along the specimen could then be

calculated (by integrating equation 6.7) and, consequently, the tensile force

distribution along the specimen can be calculated.

Juran and Chen (1988) proposed a theoretical load transfer model

derived from the (t-z) method commonly used in design of friction piles along

with the appropriate constitutive equations. Figure (6.4) shows the numerical

simulation of the pull-out resistance using this model along with the

experimental results.

In this chapter; the displacement measurements at different points along

the specimens in the pull-out tests along with the appropriate interaction laws

(relating the shear strain at any point along the reinforcement to the mobilized

shear stress at the interface) are used to develop a load transfer model. In this


135model; the interpretation procedure for the displacement-rate controlled pull-out

tests on the instrumented geogrid permits the determination of: (a) the in-soil

confined material properties, (b) the soil-geogrid interface properties, and (c) the

scale effect of the specimen length on the pull-out interface parameters. This

procedure is evaluated through comparison of the interface properties derived

from the pull-out tests with those obtained directly from the direct shear tests.

The load controlled pull-out test results are also used to asses the critical creep

pull-out load and to predict the long term creep displacement under a specified

pull-out load.


Pul

l-out

Re

sist

ance

(L

b/in

)136

Applied Load P (Lb)

0.5566

c•H

5104504J(D

EC_o

394339

H-0)a 265177124 Magnet Station No.

54321

Figure 6.1 Deformation along the Geotextile.(after Tzong and Cheng-Kuang, 1987)

150 * Unconfined Confined

125

100S ection I I __

75

Section I I I

0 10 20 30 40 50

Length of 6eotextlle (in)

Figure 6 .2 Distribution of Pull-out Resistance along Geotextile(after Bonczkewitz et al., 1988)


137

40302010Displacem ent (mm)

Figure 6.3 Applied Load versus Front and Rear Displacement (after Leshchinsky and Field, 1987)

0 .0 7

M o d e l Predict ionm 0 .0 6E»parim cntal ( Juron I9 8 5 )— 0 .0 5

U.0 .0 3 ES ■ 5B kg

Length • 55 cm0.02

CL

Front Displacem ent (mm)

Figure 6.4 Numerical Simulation of Pull-out Test on Woven Strips (after Juran and Chen, 1988)


1386.2) MODELING LOAD TRANSFER IN DISPLACEMENT-RATE CONTROLLED

PULL-OUT TESTS:

The proposed data analysis procedure is illustrated through the

interpretation of the displacement-rate controlled pull-out tests results

presented in the previous chapters. This procedure consists of the following

steps:

(1) From the Time-Nodal displacements relationship (see Figures 4.13

and 4.14), the displacements distribution along the reinforcement can be

plotted for different pull-out loads. Figures (6.5) and (6.6) show the

displacement distribution along the specimen length obtained from Tests on

geogrid 'Tensar' under confining pressure of 7 psi. Figure (6.7) shows the

displacements distribution for geogrid 'Conwed'. From the discretized elements

along the geogrid reinforcement (Figure 6.8), it can be seen that the

displacement of each element i results from the elongation of the element ( i.e.

6M - <5j) and its shear displacement 6,.

2) The strain €{ of each element i can be calculated from the relation:

€; = [ <5m - 6, ] I A x (6.8)

where and 6X are, respectively, the displacements at nodes i-1 and i, and

A x is the length of element i.

(3) The in-soil geogrid stress (force/unit width) versus strain can now be

plotted for the front element of the geogrid. Figure (6.9) shows the confined

stress-strain relationship for geogrid 'Tensar'; while Figure (6.10) shows the

same relationship for geogrid 'Conwed'. For the sake of comparison, the


139

unconfined stress-strain relationships for the two geogrids are plotted in their

respective figures. The figures show that confinement results in a substantial

increase of the stiffness modulus E.t (where t is the reinforcement thickness).

This comparison illustrates that the design criteria for geogrid reinforcement

should be based upon the confined properties derived from pull-out tests rather

than the unconfined material properties.

(4) The confined stress-strain properties (e.g. stiffness modulus E.t)

obtained under a specific confining pressure are assumed to be unique for each

geogrid material. The E.t values can be determined from the slope of the stress-

strain curve at each loading level.

(5) The tension force Tj at each node i can be calculated from the

relationship:

[ Tm - T, ] / b = E.t. [ - 6, ] / A x ........................(6.9)

(6) The shear stress distribution along the specimen can be calculated

from the static equilibrium equation of each element i:

[ T m - T J = r , . P . A x (6.10)

where P is the geogrid perimeter [= 2b].

(7) The shear stress-strain relationship at the interface can then be

plotted for all the elements of the geogrid. Figures (6.11) and (6.12) show the

shear stress-displacement curves derived from the pull-out test results on the

'Tensar' geogrid; while Figure (6.13) shows the same relationship for geogrid

'Conwed'.


140

In order to validate the analysis procedure, the shear stress-displacement

curves calculated for the 'Tensar' geogrid were compared with those obtained

directly from tests performed in the large direct shear box under the same

confining pressure and soil relative density. The results, shown in Figure (6.14),

confirm the applicability of the proposed procedure.

The analysis of pull-out test results on the geogrid reinforcement

illustrates the non-uniformity of the displacement distribution along the geogrid

length (see Figures 6.5 to 6.7). The peak shear stress mobilized at the interface

is a function of the sample length. Therefore, the extrapolation of the pull-out

test results to reinforcements of different lengths should be considered when

pull-out test results are implemented into design. The established interface

properties and confined stress-strain relationships of the reinforcement could

be implemented to back calculate the displacement distribution along the

geogrid of any length under the same testing conditions. Thereby, laboratory

pull-out test results could be extrapolated to full scsle reinforcement length.

The interpretation procedure to back calculate the displacement distribution and

the mobilized pull-out force along reinforcement of a specified length is as

follows:

(1) The reinforcement length is discretized to n number of elements, see

Figure (6.8). The displacement 6 n+1 at the end is assumed to be zero

and, consequently, the mobilized force at the end Tn+1 is also zero.


141

(2) The value of displacement at node (n) is assumed and the strain at

element n can be calculated from equation (6.8).

(3) From the established shear stress-strain relationship (Figures 6.11 to

6.13), the shear stress at the interface of element n can be obtained.

(4) The pull-out force Tn is calculated from the equilibrium equation of

element n (equation 6.10).

(5) From the confined stress-strain relationship (Figures 6.9 and 6.10), the

stiffness modulus E.t (at the strain level of element n) is obtained.

(6) The pull-out force Tn can then be calculated from equation (6.9).

(7) The value of Tn from step (6) is checked with the one calculated in

step (5). An iteration technique is used to obtain Tn value within a

limited error.

(8) The same procedure is repeated at element n-1 and up to the facing

element.

This analytical procedure was used to predict the displacements and the

pull-out forces at the nodes of geogrid reinforcements of lengths 1.5 ft, 3 ft

and 6 ft. The calculated displacements for the reinforcements of lengths 1.5

ft and 3 ft were compared, in Figure (6.15), with the measured ones. The

figure also shows the predicted displacement distribution along geogrid

reinforcement of 6 ft length. The analytical results of the pull-out forces along

the geogrids are shown in Figure (6.16). The comparison between the analytical

and experimental results illustrates the applicability of the proposed procedure.


Dis

plac

emen

t [m

m]

1421 0 0

Test—A5

80

^ BOO sec

60

dx=

400 sec [at peak]

200 sec20100 sec

50 sec

20 sec

52 3 40 1

Geogrid Nodes

Figure 6.5 Displacement Distribution Along Geogrid 'Tensar'.


143

40

Test: A -8

Geogrid Tensar SR2

30

3410. Pull-out Force [Lb]

20

CL

2680:;

10 -

2351172

2080

1140. n

Nodal Points

Figure 6.6 Displacement Distribution Along Geogrid 'Tensar'.


144

30

Geogrid: Conwed

Test: C O N -125

Pull-out at 500. sec

00>CL

80

60.

0 1 2 3Nodal Points

Figure 6.7 Displacement Distribution Along Geogrid 'Conwed'.


T .0

T GOi i+1 n n + 1

0 T = ( n+1

T. Ti+1

Figure 6.8 Numerical Analysis of Pull-out Tests.


75

60

(DL 45 o

30

15

o L

G e o g r id : T E N S A R S R 2

Confining pressure = 46 KN/m'


0 1 2 3 4 5Strain d y /d x [5tf]

Figure 6.9 Confined and Unconfined Stress-Strain Relationship for Geogrid 'Tensar-SR2'.

G e o g r id C ONW ED

C o n f in in g p re s s u r e =4B KN/m

U n c o n f in e d E x te n s io n T e s t

[6 / in height x 7 in length]

0 5 10 15 20Strain d y /d x I5U]

Figure 6.10 Confined and Unconfined Stress-Strain Relationship For Geogrid 'Conwed'.


Element # 1

Element #2

Element #3

Element #4

0 5 10 15 20Shear Displacement (mm)

Figure 6.11 interface Shear Stress-Displacement Relationship for Geogrid 'Tensar-SR2'.


148

56

r\J 42 E \Z

COCO<u

jr 20 in

oa>_cin

14

Test—A8

a•*

□ 8 a

• A

' * / "A °

+• a

a▲

▲

*

**

* A * « * ^□

• Element #1

□ Element #2

A Element #3

# Element #4

_______ I_______5 10 15 20

S h ear D isplacem ent [mm]

Figure 6.12 Interface Shear Stress-Displacement Relationship for Geogrid 'Tensar-SR2\


70

_ 56(VIE\Z

42wina)i-tno 28vs:tn

14

00 5 10 15 20

Shear Displacement (mm)

Figure 6.13 Interface Shear Stress-Displacement Relationship for Geogrid 'Conwed'.


Geogrid; Conwed

TEST: C on-1

• / • A

/

¥

~ P ar£

-/E lem en t-1

Elem ent-2

Element—3

150

5E

— 42N

□o

COCO0}]>_

00

o_cin

28

AA

O□ A

14

□ O Analytical

A Direst Shear Test

JL

0 5 10 15 20Shear Displacement (mm)

Figure 6.14 Analytical and Experimental Interface Shear Stress-Displacement Relationship for Geogrid 'Tensar-SR2\


40

EE 3 0

C0)E©© 20 _DQ.10

D 10 T>O

i K. 'a--JL.

Geogrid: Tensar SR2

Confining Pressure = 7. psi

Measured

.Analytical

4 6 B

Nodal Points10 12

Figure 6.15 Analytical and Experimental Displacement Distribution along the Geogrid Length.

4\inCL5

3TDDO

• 5 20

1 i

Geogrid: Tensar SR2

Confining Pressure = 7. psi

'••A

'•..A

'••■4

'•A .

1. fth --------------H

'■■A._

r

............................. .. i

"A ..

A " .......A ........

....... .. 1 * ...... A .....0 2 4 6 8 10 12

Nodal Points

Figure 6.16 Analytical Pull-out Load Distribution along The Geogrid Length.


1526.3) ANALYSIS OF LOAD-CONTROLLED PULL-OUT TESTS:

The process of time-dependent deformation of the geogrid subjected to

a constant stress displays generally three stages; in addition to its

instantaneous deformation. These stages are (see Figure 4.18):

i) Stage I (Primary Creep): A non-steady state creep where deformations

develop at a decreasing rate.

ii) Stage II (Secondary Creep): A stage of steady state creep where the

deformation rate reaches a minimum value and then stays constant.

iii) Stage III (Tertiary creep): Where the deformation develops at an

increasing rate and ends at failure.

The duration and effect of any particular stage vary with the geogrid type and

load. The engineering concern in the long term stability of reinforced earth

structures is to determine the critical load level and displacement at which the

secondary creep stabilizes and before the tertiary creep starts to develop. Since

these properties depend on the time available for observations, an accurate

experimental approach is a difficult and time consuming task. Deformations

which may appear to be stabilizing may develop if observed for a longer period

of time.

The instrumentation for displacement measurements along the geogrids

permitted the calculation of creep strains when the geogrid is subjected to

stepped creep loads. Pull-out tests were conducted on 'Conwed' geogrids

where the specimens were subjected to creep loads of 50%, 65% and 85% of


153maximum pull-out load. Figures (6.17) and (6.18) show the displacement

measurement at different nodes along the geogrid at 65% Tmax and 85%

Tmax, respectively. The calculated creep strains are plotted with time in Figures

(6.19) and (6.20) for load levels 65% Tmax and 85% Tmax, respectively.

Figure (6.20) shows accelerating creep strain at the "Conwed' geogrid when

it is subjected to pull-out load of 85% Tmax. The unconfined creep strains

under the same loading levels are plotted in their respective figures. The

comparison between the unconfined and confined creep strains shows that

confinement causes lower creep strains under the same loading levels.

The creep strain under a sustained load at time t can be analyzed by

utilizing the pull-out creep test results into the relationships proposed by Singh

and Mitchell (1968). These relationships are based on the following

observations:

(i) A relationship between strain rate ie) and time exists for many clays

subjected to loadings within the range of 30% to 90% of strength. This

relationship is illustrated by a form of a plot between the log-strain rate and log

time (Figure 6.21). The Figure shows that log-strain rate decreases linearly with

log-time with a slope (m) independent of creep load. This relationship can be

expressed as:

Log € = Log e(t,,T) - m Log ( t / t , ) ....................................................... (6.11)

Where, e = strain rate at any time t,

€ (t,,T) = strain rate at unit time t, and load T.


154

(ii) A relationship between log-strain rate and the practical range of

applied loads (T) can be expressed linearly (Figure 6.22) as:

Log e = Log e(t,T0) + a T (6.12)

Where, e(t,T0) = strain rate obtained by projecting the straight line of the

relationship to T = 0,

a = slope of the straight line portion of the log-strain rate versus load.

Combining equations (6.11) and (6.12) and integrating we get:

A a l 1 -me = €y + ( ------------) e (t - 1) for m ^ 1 .....................(6.13-a)

1-m

aT€ = ey + A e Log (t) for m = 1 (6.13-b)

Where A = e(tv T0). Equation (6.13) is a simple three parameter relationship

between strain and time which is found to be applicable to different kinds of

soils (Singh and Mitchell, 1968). Shrestha and Bell (1982-b) investigated its

applicability to predict creep strains for different kinds of geotextiles. The

relationship between log-strain rate versus log-time for geogrid 'Conwed' is

shown in Figure (6.23) for different pull-out levels. The figure shows a linear

relationship at load levels of 50% and 65% of Tmax. Figure (6.24) shows the

relationship between the log-strain rate versus loading levels T normalized with

respect to the maximum pull-out load Tmax.


Nod

al

Dis

plac

emen

t (m

m)

10

TEST: C 0 N - C 3

C o n fin in g "P re s s u re 14 psi

L o a d : 6 5 % T m a x

Node #0

2 -

tyodes ^ 3 .4

0 20 40 60 80 100 120Time (Hour)

Figure 6.17 Nodal Displacements at Load-Controlled Pull-out Teston Geogrid 'Conwed'

sBaQ6©Oe1*£3

20

TEST: C0N-C3

C onfin ing Pressure 14 psi

Load 85% Tmax Node #015

10

Node #15

080 12020 40 60 1000

Time (Hour)

Figure 6.18 Nodal Displacements at Load-Controlled Pull-out Teston Geogrid 'Conwed'


15610

(3.0 Kips)

Unconfined Extensiona

Load (2.4 Kips)

a

4 Confining Pressure • 14 psi

65% Tnax (2.8 Kips)

2

Geogrid Conned

o1000100101

Figure 7.19 Confined & Unconfined Creep Strain for Geogrid 'Conwed'.

IS

Load (3.7 Kips)

Unconflned Extension

10

Pressure 14 psiConfining

BSX T u x (3.6 Kips)

s

6eogrld Conwed

o 1000100(H oar)

Figure 7.20 Confined & Unconfined Creep Strain for Geogrid 'Conwed'.


157

0 .002

0.0016I4> .0004<0cr

.0 0 0 2c(0L. .0001

.00004Oioko AIKiviol Cloy Wolir Conltnl«63% (AMir MyfOfOfno ond Shibolo,t953l

.000010.1 2 4 10 .20 40 too

Tine- Bin

Figure 6.21 Strain-Rate versus Time Relationship. (After Singh and Mitchell, 1968)

I0.0

Rtmoldfd SonFioncilceBoj Mud Wolif Conltni S2*A

2 .0

1.0

0.4 05 ----

0.2«yAi10ccc

L.i/i

.0 2Ii

.004

•002

0.19 0.20 0.21 0.22 0.23 - 0.24 ' 0.23 0.23

O evlatorlc Streas

Figure 6.22 Strain-Rate versus Stress Relationship. (After Singh and Mitchell, 1968)


158

0.050.03

50% Tmax0.02

□A 85% TmaxA A.0.01

o n

0.005

0.002Geogrld: Conwed

0.001 20 30 50 100 200 300 500 1,000Time (m in)

Figure 6.23 Strain-rate versus Time Relationship for Geogrid 'Conwed'.

0.1•R= 20 min

>nt= 40 min0.05

0.03

0.02t= 120 min

0.01on

0.005

0.002

Geogrid Conwed0.001

0.4 0.5 0.8 0.9 1Load (T/Tmax)

Figure 6.24 Strain-rate versus Load Relationship for Geogrid 'Conwed'.


CHAPTER 7

STABILITY ANALYSIS OF REINFORCED SOIL STRUCTURES

7.1) INTRODUCTION:

The stability analysis of reinforced soil structures involves the evaluation

of the following:

(i) The external stability when the potential sliding surface develops

outside the reinforced zone. In this case, the conventional slope stability

procedures like modified Bishop's analysis can be used.

(ii) The internal stability of the reinforced structure when the failure

surface passes through the reinforced zone. In this case, the mobilized tensile

forces in the reinforcement along with the soil shear stress along the failure

surface resist the driving forces induced by the weight of soil in the active zone

(Figure 7.1-a).

The internal stability of the soil-reinforcement system requires the analysis of

two failure criteria:

(i) Failure by pull-out of the reinforcement: the mobilized tensile forces of

the reinforcement should not exceed its pull-out resistance.

(ii) Failure by breakage of the reinforcement: the mobilized tensile force in

reinforcement should not exceed its ultimate tensile resistance.

Most of the internal stability analysis of the reinforced structures are

159


based on limit equilibrium methods which are modified to account for the

reinforcement tensile resistance. These methods consider different

assumptions with regard to the failure mechanism (rotational or translational),

the shape of the failure surface (plane, bilinear, circular and log spiral) as well

as the deformation and orientation of the reinforcement. Rowe (1984) and Hird

(1986) used the limit equilibrium design method with circular sliding surface to

obtain design charts for reinforced embankments on soft foundation soils.

Milligan and La Rochelle (1984) developed limit equilibrium design charts with

both a circular arc for rotational failure and a planner sliding surface for

translational failure of embankments over soft foundations. Murray (1982)

considered a bilinear slip surface to evaluate the safety factors of reinforced

slopes. Jewell etal. (1984) and Schmertmann etal. (1987) presented a bilinear

wedge limit analysis method to evaluate the rotational and translational failure

mechanisms in steep reinforced slopes. Leshchinsky etal. (1986), Leshchinsky

and Perry (1989) and Leshchinsky and Boedeker (1989) presented an analytical

approach based on limit equilibrium analysis of log spiral slip surface and using

variational calculus to find the rotational and translational failures in reinforced

slopes.

Since the limit equilibrium analysis cannot predict the actual distribution

of the tension force at the failure state, different assumptions regarding the

distribution of the tensile force T were considered. In these methods the

tension forces are usually assumed to follow either the K0 or the K. Rankine's


161

earth pressure distribution. Leshchinsky and Perry (1989) considered Tf as a

linear function of the overburden pressure; i.e:

Tj = T , (H - yj) /H

where yj is the is the depth of the reinforcement j. Christopher et al. (1989)

considered a distribution factor (K/K.). This ratio depends on the stiffness of the

reinforced mass Sr which is primarily related to the reinforcement extensibility

and its horizontal and vertical spacings in the soil mass.

Different assumptions regarding the orientation of the reinforcement at

the failure surface were also considered. As the tension forces are mobilized,

the reinforcement reorients itself from its horizontal position {0 = 0) to an angle

0 (Figure 7.1 -a). The orientation of the reinforcement {0) governs the magnitude

of the mobilized tension forces and consequently the factor of safety. Jewell

(1982) and Ingold (1982) assumed that the tension forces stay horizontal;

while Brakel et al. (1982) assumed the direction of the tension forces to be

tangent to the failure surface. Gourc et al. (1986) considered both limit cases

assuming horizontal forces at small displacements; whereas at large

deformations, the reinforcement becomes tangential to the potential failure

surface. Rowe and Soderman (1985) performed finite element study to evaluate

the effect of reinforcement orientation on the moment arm of the failure circle

in reinforced embankments. They concluded that the moment arm for

calculating the resisting moment due to geotextile force can be taken as

moment arm = (R + z)/2 for D/B <; 0.42


162

moment arm = z for D/B > 0.42

Where R, z, D and B are as shown in Figure (7.1-b). Leshchinsky and Perry

(1989) determined that a value of /? equals (0-0) is required to mobilize the

reinforcement maximum tensile force where 0 is the inclination of the failure

surface with horizontal and 0 is the friction angle of the soil.

The slope stability methods based on limit equilibrium can be used to

evaluate the safety factors with respect to the failure mechanism along the

assumed failure surfaces. Nevertheless, these methods can not provide an

estimation of the maximum tension forces that would be generated at each

reinforcement level. Consequently, they do not allow for the evaluation of the

local stability of the reinforced soil mass at the level of each reinforcement with

respect to the potential failure due to breakage or pull-out of the reinforcement.

Moreover, these methods do not allow for an evaluation of the effect of soil

dilatancy and reinforcement extensibility on the generated tension forces and

structure stability. The development of a working stress analysis for the

evaluation of the tensile forces of the reinforcement and the stability of the

structure requires adequate constitutive equations for the soil, in-soil confined

stress-strain properties of the reinforcement and an assumption pertaining the

strain compatibility between the soil and the reinforcement.

This chapter presents a soil-reinforcement analytical model that

associates an elasto-plastic strain hardening soil behavior, elastic-perfectly

plastic response of the reinforcement and a perfect adherence at the soil-


163

reinforcement interface to calculate the maximum tension forces in the

reinforcement. This model (Juran et al., 1988-b,c and 1990) allows for the

evaluation of the effect of soil dilatancy and reinforcement extensibility on the

generated tension forces.

7.2) MODELING GEOSYNTHETIC-REINFORCED SOIL WALLS:

The model used in the analysis of geosynthetic reinforced soils is based

on the analogy between the response of the reinforced soil material to shearing

(Figure 7.2-b) and the plane strain shear mechanism developed along the

potential failure surface (Figure 7.2-a). This analogy has been assessed through

numerical simulations presented by Jewell (1980) and Juran et al. (1988-b and

c). The main design assumptions of this model are:

(i) The soil is assumed to be homogeneous, isotropic and follows an

elasto-plastic strain hardening (and strain softening) yield function with a non

associated flow rule.

(i) An active zone develops in the soil which undergoes a rigid body

movement along a plane failure surface (Figure 7.2-a).

(ii) The shear displacement of soil along the potential failure surface is

assumed linear with depth and the mobilized extension of the reinforcement

along the shear zone satisfies the strain compatibility with the specified soil

shear strain.

(iii) The reinforcement is assumed elastic-perfectly plastic and the generated


164

tension force is retained by the portions of the reinforcement in the resisting

zone.

The soil-geosynthetic model is based on the development of the following

relationships: (i) Constitutive equations of the soil; (ii) stress-strain relationship

for the reinforcement; (iii) soil-reinforcement interaction; and (iv) equilibrium

conditions in the soil mass.

i) Soil Model:

The yield surface of soil is defined by a yield function that obeys Mohr-

Coulomb's failure criteria. It is assumed that the soil layer along the potential

sliding surface undergoes simple shearing and therefore the strain hardening is

a function of the shear strain at the failure surface yxy. The yield function can

then be written as :

T xy----------------h ^ ) = 0 ( 7 .1 )

ay

where and oy are the mobilized shear and normal stresses along the

potential sliding surface, respectively; and hty^) is the strain hardening function

and equals tan Qm where <pm is the mobilized soil friction angle at the failure

plane.

For dense dilating sand, it is assumed that the strain hardening function

hfKxv) is parabolic (Figure 7.3) and is written as:

c K*y (Kxy - a)= .......................... (7 .2 )

( K x y + b ) 2


165

Where the constants a, b and c are determined from the initial conditions of the

strain hardening function which yield:

a = - 4 (ay/G) ( tan2 0 P . L 2 ) / tan 0 „

b = 2 (<7y/G) ( tan 0 P . L )

c = tan

and, L = 1 + [ 1 - (tan 0 ^ / tan 0 P)]*

Where G is the initial soil shear modulus and 0P and 0 „ are, respectively, the

soil friction angles at peak and at critical state (Figure 7.3). The determination

of these constants along with the detailed mathematical formulation of the

equations presented herein are shown in Appendix D.

The plastic flow of the sand is defined by its dilatancy angle u which

equals:(3Cy 1 Tjy

t a n u = = — [ t a n 0 CV - — ] ( 7 . 3 )dKxy M

where dey is soil extension (or compression) increment in the direction of

orthogonal to the failure surface, dKxy is the shear strain increment along the

failure surface and // is a dilatancy parameter equals f/, when 0 m 0 CV and a

contracting parameter equals jj2 f ° r 0m > 0cv

(ii) Reinforcement:

The reinforcement is assumed to be elastic-perfectly plastic material. The

maximum tensile force of the reinforcement should not exceed either its elastic

limit determined from confined tests on the reinforcement under the specified


166

confining pressures or its pull-out resistance.

(iii) Soil-Reinforcement Interaction:

As the tension forces are being mobilized, the reinforcement reorients

itself from its original horizontal position to an angle /? (Figure 7.2-a). The

current inclination of the reinforcement.^ is calculated incrementally from Figure

(7.4-a) in the equation:

tan A„, + dKxytan A, = ........................... (7.4)

1 + deyn

with an initial horizontal inclination Aa = — - a ;where a is the slope of the2

failure surface.

The soil deformation in the direction of the reinforcement At for each yKy

at the level of each reinforcement can be calculated from the Mohr's circle of

strain (Figure 7.4-b). The strain in soil 6ep is given by:

co s (A -u ) s i n A,d e , = t a n i> [ 1 ------------------------------------ ] d7 ) [ y ............................. ( 7 . 5 )

s i n v

The assumption of a perfect soil-reinforcement adherence implies that the

tension increment of the reinforcement deR is equal to that of the soil in its

direction (i.e. de,). The elastic-perfectly plastic constitutive equation of the

reinforcement enables to calculate the tension stress increment da R :

doR = E . deR , with oR ^ RT (7.6)

where E is the reinforcement elastic modulus and RT is its yield stress. It should

be noted that a non-linear elastic behavior of the reinforcement can be


167

numerically implemented in the model. The maximum tension force in the

reinforcement per unit area along the failure surface is then given by:

T,n« E.t.b = f R ........... (7.7)Sh- Sv Sh. Sv

where t and b are the thickness and width od the reinforcement, respectively;

and Sh and Sv are the horizontal and vertical spacings between the

reinforcements.

(iv) Equilibrium Condition:

The forces acting on the active zone are shown in Figure (7.2-a). The

gravity force in the active zone is resisted by the tension forces in the

reinforcement and the soil reaction along the potential sliding surface. The

gravity force in the active zone is given by :

1W = 'A p H02 [ ------------ tan J ] (7.8)

tan a

where p is the soil unit weight and J is the inclination of the wall facing with

respect to vertical. The analysis of equilibrium condition in the active zone

yields a nondimentional solution, relating the actual structure height H0 to the

specified shear strain level y*y along the potential sliding surface :

E.t.b 2 t ZN f R sinW +o) cot(a-Q) - XN eR cosM + o)]H0 / [ -------------] = ---------------------------------------------------------------------- . . (7.9)

p.Sh.Sv N [ (1/tan a) - tan J ]

The height of the structure H0 is minimized with respect to a in equation (7.9)

in order to derive the most critical solution, i.e. 3H0/9o = 0 .


168

FailureSurface

Active Zonemax

Resistant Zone

X

Figure 7 .1 -a Schematic Diagram of Failure Mechanism in Reinforced Soil Walls (After Gourc et al. 1986).

Fill

Clayey Soil

; Geotextile

Figure 7.1 -b Schematic Diagram of Failure Mechanism in Reinforced Embankments (After Rowe and Soderman, 1985).


169

Tmox

THIN LAYER OF SHEARED ZONE

REINFORCEMENT

REINFORCED SOIL WALL DIRECT SHEAR ANALOGY

Figure 7.2 Analogy between Behavior of Reinforced Soils in Retaining Wall and in Direct Shear Test (After Juran et al. 1990)

to n <f>

G/cr,<c

>»

Figure 7.3 Elasto-plastic Strain Hardening of Soil.


170

DIRECTION OF REINFORCEMENT

d * d<d «

T 1 H I IZERO EXTENSION LINE (Horlionlol)

(a) Deformation of Reinforcement (b) Mohr's Circle of Strain

Figure 7 .4 Modeling Behavior of Reinforced Soil. (After Juran et al. 1990)


171

7.3) ANALYSIS OF BEHAVIOR OF GEOSYNTHETIC SOIL WALLS:

The formulations of the relationships of the proposed strain compatibility

analysis are presented in detail in Appendix D. These relationships were

implemented in a computer program to predict for each structure height the

maximum tension forces mobilized in the reinforcements and the location of the

potential failure plane. Figure (7.5) shows a flow chart illustrating the

computational scheme. The output data can then be used to evaluate the

potential failure of each reinforcement due to either breakage or sliding. The

selection of the appropriate reinforcement and spacings should satisfy the

required safety factors with respect to these potential failure modes, as well as

design criteria pertaining the admissible strains in the reinforcements.

The maximum tensile force in the reinforcement depends on the stiffness

of the reinforced soil mass which is primarily related to the extensibility and

density of the reinforcement. The effect of the relative stiffness of the

reinforcement on the maximum tension forces is illustrated in Figure (7.6). The

relative stiffness of the reinforcement is defined as:

h0 = E .t.b /p .S h.Sv

The maximum tension force is given by the non-dimensional parameter:

= fm®/ P-H.Sh.Sv

where H is the height of the structure. For a given soil type, a unique TN =

f(H/h0) relationship can be developed provided that the reinforcement yield

strain is not exceeded (i.e. eR < £y). The relationship in Figure (7.6) was


172

calculated for a typical range of compacted backfill materials with the following

soil parameters: 0P = 40°, <pcv = 31°, p = 15.5 KN/m3anda constant dilatancy

angle = 1/3 <pp = 13°, and a range of normalized shear modulus (G/av). The

figure illustrates several aspects of the behavior of reinforced soil wall:

1- For a given structure height, the relative stiffness, h0, of the

reinforcement significantly affects the state of stress in the soil. As h0

decreases (i.e. extensibility increases), the normalized tension forces, TN,

decreases from Kc state of stress (for quasi-inextensible reinforcement) to K.

state of stress (for highly extensible reinforcement).

2- the variation of the normalized tension forces, TN, with the relative

stiffness, h0, depends upon the normalized shear modulus, G/oy, which governs

the shear strain in the soil.

In order to assess the effect of soil dilatancy on the locus of maximum

tension forces, two assumptions were considered: (1) constant dilatancy angle

= 13°, and (2) variable, strain dependent dilatancy angle defined as in Equation

(7.3) with the contractancy parameter//, = 10 and the dilatancy parameterp 2

= 0.9, for the sand used in this parametric study. It should be indicated that

the assumption of variable dilatancy attempts to simulate soil volume change

with an initial phase of contraction followed by a phase of dilation. Soil

dilatancy is highly dependent upon the in-situ confining pressure and

construction process. The two assumptions specified above yield the probable

range of the actual soil behavior in a compacted backfill material. The effect of


173soil dilatancy on the relationship between relative stiffness of the

reinforcement, h0, and the normalized tension forces is shown in Figure (7.7).

The Figure shows that the assumption of a constant dilatancy rate (i.e., um =

u ^ ) leads to more conservative results.

The effect of the reinforcement relative stiffness ,h0, on the inclination

of the failure surface is shown in Figure (7.8). The Figure illustrates the

following:

1- For a given structure height, as the relative stiffness, h0, increases (

(as for quasi-inextensible reinforcement), the inclination of the failure surface

increases and becomes practically vertical.

2- As the relative stiffness decreases (e.g. higher extensibility of the

reinforcement), the inclination of the failure progressively approaches that of

Mohr-Coulomb's failure plane (i.e. a„ = nl4 + 0/2). This variation of aa is

consistent with observations on both reduced scale and centrifugal model walls

(Juran and Christopher 1989; and Jaber et al. 1987).

These results also conform with observations on walls reinforced with

both extensible and quasi-inextensible reinforcement (Juran and Chen 1989;

and Christopher et al. 1989). These observations are illustrated in Figure (7.9)

for both types of reinforcement. The Figure shows practically vertical failure

surfaces at the upper part of the structure when inextensible reinforcement is

used. It should be noted that the vertical failure surface, when inextensible

reinforcements are used, suggests a probable more critical failure height when


174

a log spiral failure surface is assumed.

The design charts presented in Figures (7.6) to (7.8) can be used to

evaluate the local stability of each reinforcement and to predict the failure

surface for the specified soil data. The analysis allows the engineer to design

and analyze the structure behavior under the design working load. In order to

evaluate the design assumptions considered in this analysis, the predicted

tension forces are compared with those measured in model wails and

embankments reinforced with different types of geosynthetics.


175Input design data:

Soil :p,<pp, 0 CV, G/cry, p, Reinf.: E, t, b, Rt

Wall: J, Sh, Sv

1 t

Calculate max. shear strain:Kj+i = Kj + dKxv

1l .Assume inclination a *

r

Assume linear shear strain distrib. Calculate for i (i = 1,2,..N):

0mii) Eqn. (7.2) tU i Eqn. (7.3)

Eqn. (7.5)

r

Calculate for each reinf. i (i= 1,..N) f R0J, Eqn. (7.6)

Tjj < Rt Eqn. (7.7)

t ...

Analyze equilibrium at active zone Hoj Eqn. (7.9)

No

Figure 7 .5 Flow Chart of Computational Scheme.


0.5Vertical Wall

JZ 0.4Ko G/cr = 30

G/o- = 60

G/cr = 120

</>

0.2Ka

o.i 0.30.250.20.150.10.050Normalized Height (H /h o )

Figure 7.6 Effect of Reinforcement Stiffness, ha, on MaximumTension Forces.

0.5

Vertical Wall

<t> = 4 o' G /c r = 60SZ 0.4 Ko

Q- 0.3

0.2

0.1 0.15 0.2Normalized Height (H /h o )

0.25 0.30.05

Figure 7.7 Effect of Soil Dilatancy on Maximum Tension Forces.


17790

Vertical Wall 4> - 40 °B5

© 80 O)

75

U. 70

G/cr — 30 G/cr = 60 G /cr= 120

65

60 0.20.05 , 0.1 0.15Normalized Height (H /h o )

o

Figure 7.8 Effect of Reinforcement Stiffness, h„, on the Inclinationof Failure Surface.

0 3 H |

(b) inextensible reinforcements

H - lo nU.------- T . |

(c) extensible reinforcements

Figure 7.9 Observations on Inclination of Failure Surface for Different Reinforcements (Christopher et al. 1989).


178

7.4) ANALYSIS OF MODEL WALL RESULTS:

The tension forces predicted using the proposed analysis were compared

with those measured in different model walls and embankments, namely:

1- Two laboratory model test walls, of 60 cm high, reinforced with non

woven geotextile strips (Juran and Christopher, 1989).

2- Three instrumented full scale reinforced soil walls of 20 ft high

(Christopher et al., 1989).

3- Two instrumented embankments of 20 ft high and slope 0.5 horizontal

to 1 vertical and two embankments of 25 ft high and slope 1 horizontal to 1

vertical (Christopher et al., 1989).

(i) Analysis of Laboratory Model Tests:

Figure (7.10) shows typical instrumented cross section of the model

walls and the observed failure surface. The material properties of the non

woven geotextile for both walls is shown in table (7.1) and the confined-

unconfined extension properties of the reinforcement are shown in Figure

(7.11). For the purpose of the analysis, the reinforcement is assumed to be

linearly elastic perfectly plastic material. The soil used in the study was

Fontainbleau sand, compacted to an average dry density of 15.5 KN/m3, with

the following properties: 0 P = 40°, 0 CV = 31.5°, and Gla = 50. For the sake

of analysis, the two assumptions of constant and variable dilatancy angles were

considered.


Figure (7.12) and (7.13) show the predicted and measured distributions

of the tension forces in walls No. 1 and No. 2, respectively. The measured

tension forces in the model walls are close to those predicted by K0 state of

stress at the upper part of the walls and approach those predicted by K, state

of stress (i.e. Rankine's state of stress) at the lower part of the walls. The

numerical analysis of the model walls, assuming a constant dilatancy rate,

predicts fairly well the measured tension forces in wall No. 1 while it

overestimates the tension forces in wall No. 2. The tension forces in the

reinforcement are significantly affected by soil dilatancy which is difficult to

assess. However, the assumption of a constant dilatancy rate appears to

provide a conservative bound for the design values of tension forces in the

reinforcement. For both models, the numerical simulation shows that as the

lower reinforcements attain their limit tension force, the lateral thrust due to the

increase in the structure height is progressively transferred to the upper

reinforcement.

Wall Reinforcement Elastic LimitFv/b (N/cm)

E.t(N/cm)

b(cm)

sh(cm)

3 *(cm)

L(cm)

Failure

1 Non-woven geotextile 6 500 6 30 7 60 Breakage

2 Non-woven geotextile 6 500 5 30 5 50 Breakage

Table 7.1 Material Properties and Geometrical Parameters of Test Walls(Juran and Christopher, 1989).


H (cm)

• Right Facing o Left Facing

a Breokage in the Strips

1 2 3 4 5POTENTIOMETER

INSTRUMENTATION SCHEME FAILURE SURFACE

Figure 7.10 Geometry of Test Walls and Failure Surfaces.

t - ( N /c m )

CONFINED

ES/b = 6 9 N/cm

ES/b = 8 .3 N/cmUNCONFINEDE S /b = 5 0 0 N/cm

ES/b=28,6 N/cm

Figure 7.11 'Confined' and 'Unconfined' Extension Tests on Non-woven Geotextile (Juran and Christopher, 1989).

180


181

80

70

• Experimental

A d Calculated (v =13°)

^ Calculated {v variable)60

O 50 - A.Strip 3

£,401

Strip BX30 Kb.

Strip 2

20

Ka

Model- 1

0 10 20 30 40 50 60Tensi le Force, T m a x [N]

Figure 7.12 Measured and Predicted Maximum Tension Forces in Wall No. 1.


80

• Experimental □ Calculated (2 = 13°)

▲ Calc, (t* variable)

70

60

50

Strip K

Strip J

a) 30 *o

Ka Ko

20 Strip H

Ka Model-2

0 10 20 30 40 50 60Tensile Force, Tmax [N]

Figure 7.13 Measured and Predicted Maximum Tension Forces in Wail No. 2.


183

(ii) Analysis of Full-Scale Model Walls:

Three full scale model walls of 20 ft height (Christopher et al. 1989)

were evaluated in this analysis. Typical geometry and soil properties of the

walls are shown in Figure (7.14). Different reinforcements were used in these

walls with properties shown in Table (7.2). The relative stiffness of the

reinforcements in these walls are defined as S, = (E.t.b)/(Sh.Sv), and it relates

to the relative stiffness parameter used in this analysis by the relationship (he

= Sr/p) where p is the soil density.

Figures (7.15) to (7.17) show the comparison between the measured

tension forces in the reinforcements and the predicted ones for the three walls.

Two different cases of soil dilatancy were used to predict the tension forces;

namely, fully mobilized soil dilatency , vm = i w and variable dilatancy angle

,vm~ f (ym) . The figures show that the model predicts fairly well the maximum

tension forces in the walls and that the assumption of fully mobilized soil

dilatancy gives an upper bound of the developed tension forces.

The measured and predicted maximum tension forces in the three walls

are normalized with respect to Rankine's state of stress, K„ in Figure (7.18).

The variation of K/K. with depth illustrates the development of the state of

stress in the reinforcement. The figure shows that the distribution of the

maximum tension forces is close to Kc state of stress at the upper part of the

wall and approaches Ka at the lower part, when the flexible reinforcements are

used in walls No. 2 and No. 7.


184

This observation is similar to the measured and predicted behavior of non

woven geotextiles in the test walls analyzed previously. When rigid reinforce

ment is used, as in wall No. 3, The distribution of the maximum tension force

remains close to K0 state of stresses.

Wall Reinforcementtype

sr(K/ft2)

sh(ft)

Sv(ft)

2 Geogrid 56 1.0 2.5

3 Bar mats 1037 4.92 2.5

7 Gabion woven mesh 20 1.0 2.5

Table (7.2) Material Properties and Geometrical Parameters of Model Walls.(Christopher et al. 1989)


185

6 Compacted SoO Layers

Backfill Uateriol : Gravelly Sand7 - .130 pcf 4 — 40 degrees c “ 0.0 p i t

Reinforced Zona

FoundoUon Soil :Uedlum Dense Sond With Grovel Traces (SP-SW)

Physical Conditions

Figure 7.14 Geometry and Soil Properties of Model Walls.


186

20

15

j rD)

•5 1° IE

AA

W all-2

a\V Geogrld Tensar SR2

Sr = 56 K lps /ft2

", \ o Measured%

\ \ • Calc. ( V = 1 2 ° )%%

'■# q a Calc. ( v var.)

0 o •..-o

\ * £IA

0.5 1 1.5 2 2.5Maximum Tension (Kips/ft)



187

20

15

JZO)*5 10X

D£

V .%• \v . *•.V.. V • V. Wall-3VSL bar mats

^ Sr = 1037 K lps /ft2

\ \ Measured% •,

\ V , Calc. ( ^ = 12°)'• % \

\ \ Calc. ( 2) var.)0 • \

IIa

■1b

\ ^<r A • \ t >

0 0.5 1 1.5 2 2.5 3Maximum Tension (KIps/ft)



188

20

\ W a ll-7\ \Gabion Woven Mesh

15 t- * \Sr = 20 K Ip s /ft2

s :D)

*5 10 X

o Measured

• Calc. ( = 12)

V A Calc. ( V var.)*

& \TIA r-T

• •

■U— l0.5 1 1.5 2 2.5

Maximum Tension (K Ips/ft)



189

20

D>

Measured Sr

(K /ft?

o Wall—7 20

A W a ll-2 56

□ W all-3 1037

2.50.5K/Ka

Figure 7.18 State of Stress in The Reinforcement with Depth.


190

(iii) Analysis of Reinforced Embankments:

The behavior of four instrumented embankments were evaluated using

the proposed analysis. The geometry and properties of the embankments are

shown in Table (7.3). In these embankments, eight layers of reinforcement with

a uniform spacing of 2.5 ft and total length of 14 ft were used. Figures (7.19)

to (7.22) show the comparison between the measured and predicted maximum

tension forces in the reinforcement. Figure (7.23) shows a comparison between

the predicted inclination of failure surface and the measured locus of maximum

tension forces in embankment No. 1. The results show that the proposed

analysis predicts fairly well the distribution of tension forces and inclination of

the failure surface in the reinforced embankments.

Embankment Slope Reinforcementtype

Tuhlb/in

E.AKips/ft

1 0.5 H: 1 V Geogrid 145. 90

2 0.5 H: 1 V Woven geotextile 214 96

3 1 H: 1 V Geogrid 145 90

4 1 H: 1 V Woven Geotextile 214 24

Table 7.3 Properties of Reinforced Embankments. (Christopher et al. 1989).


20i

V E m bankm ent-1

15

•5 10x

D£

%

Slope: 0.5 H / 1 V

V i Geogrid (EA=90 K ip s /ft)Q \I

o Measured

a Calc. ( 1 = 10°)

\ * i \ • Calc. ( %> var.)

Limit Equil.

V* A \

-L<k-0.2 0.4 0.6 0.B 1 1.2 1.4Maximum Tension (Kips/ft)

Figure 7.19 Measured & Predicted Tension Forces in Embankment 1.


20

15 I- ©

Em bankm ent-2

Slope 0.5 H /1 V

Geotextlle (EA=96 K Ips /ft)

O Measured

,#A A Calc. ( » = 1 0 ° )

x: \ \ • Calc. (*> var.)

•S* 10 ~ $ ................... ...... Limit Equil.\m aX

Q%

O

-A-L0.2 0.4 0.6 0.8 1 1.2 1.4Maximum Tension (Kips/ft)



25

20

15 «+—JZo>

*0)X= 10 o£

> Em bankm ent-3

V Slope: 1 H /1 V

' \ * Geogrid (EA=90 K Ips /ft)•i A •%

o Measured

\ \ A Calc. ( 2) = 1 0 °)*• t\ \» a • Calc. ( 2? var.)

\ \ Limit Equil.

\ PX » A

01• A

<6 A

• 1 A0.2 0.4 0.6 0.8 1 1.2

Maximum Tension (KIps/ft)



194

ZD tI;

u Embankment—4mi1 ••1 \ Slope: 1 H /1 V

201

V1 \

Geotextlle (EA=24 K Ips/ft)

N 9 \1 * • i •

o Measured4"S -/

1511

- \ i a Calc. ( * ? = 10°)JZD> i * \i • Calc. ( t> var.)'55X

i\ i \ ......... Limit Equil.

n 10 - 9 \£ V\\

•

5V

- /d i \ - '

0 ------- *4— t i i i * A _________ I_________ l_________ I_________ i0 0.2 0.4 0.6 0.8 1 1.2

Maximum Tension (Kfps/ft)



195

Embankment 1Slope = .5 H / 1 V Geogrid

o Measured ( 1 2 /4 /8 7 ) — Computed

T- 25

Embonkment Height ■ 20. ft.

20So Range of T ,,^ from Limit

Equilibrium Analysis

- 1 5-P

.P- 1 0

- 5

12 18 24 30Distonce From Embankment Toe (ft.)

36

Figure 7.23 Measured & Predicted Failure Surface in Embankment 1,


Hei

ght

(ft.

)

CHAPTER 8

CONCLUSIONS

The pull-out box and large direct shear box were designed and

constructed in order to evaluate the short term and long term interface

parameters and confined stress-strain properties of the geogrids. Displacement-

rate controlled tests were conducted to provide the parameters related to short

term performance of the geogrids; namely, their pull-out resistance, front

displacement at peak, and interface stiffness modulus. The instrumentation of

the pull-out box enabled the measurement of the nodal displacements along the

confined geogrid. These measurements were utilized in a load-transfer model

in order to determine the stress-strain properties of the geogrids. Load-

controlled pull-out tests were conducted to evaluate the long-term parameters

of the geogrids; namely, the critical creep load under specific confining

pressures, and creep displacement.

The results of the pull-out tests on both 'Tensar' and 'Conwed' geogrids

in granular soils illustrate the importance of soil density and the applied

confining pressures on the pull-out behavior. The high soil density and confining

pressure increase the geogrid frictional resistance at the interface and increase

the lateral earth resistance on the transversal elements. Consequently, they

restrain the displacement of the geogrid along its length and increase the

geogrid pull-out resistance.

196


197

Pull-out testing program was conducted to evaluate the performance of

the facility and to assess the effect of different testing parameters (i.e. testing

modes, pull-out rate, soil density, confining pressure, and box boundary

conditions) on the interface properties. The parametric study led to the

following conclusions:

(i) The increase in the pull-out displacement-rate results in a decrease in

the peak pull-out resistance and interface stiffness modulus. Pull-out tests

results appears to be less sensitive to changes in displacement-rates under 6

mm/min.

(ii) Pull-out test results on the geogrid in different soil densities illustrate

the influential effect of soil placement and compaction. A compaction

procedure was developed and calibrated to obtain uniform soil density along the

geogrid. Different soil placement techniques and compaction procedures have

yet to be developed if the geogrid is to be tested in fine grained soils.

(iii) Test results on the geogrid under different confining pressures

demonstrate that the peak pull-out strength is mobilized at strain levels much

lower than those of the unconfined extension tests. This conclusion suggests

that the design criteria of reinforced-structures should be based on the pull-out

tests results rather than the unconfined extension test results.

(iv) The lateral earth pressure developed at the facing increases the soil-

geogrid resistance to pull-out. The use of sleeves around the pull-out slot

transfers the soil-geogrid interface resistance far behind the rigid front wall,


198

which imposes a practical solution to reduce the front rigid wall effect. Pull-out

tests with different sleeve lengths led to the conclusion that a minimum sleeve

length of 1 ft is enough to eliminate this boundary effect.

(v) Another boundary condition that influence pull-out test results is the

upper and lower box boundaries. When small soil thickness is used, the rigid

top and bottom plates increase soil friction at the soil-geogrid interface and

restrain soil dilatancy. Pull-out tests on geogrids in different soil thicknesses

demonstrate that a minimum soil thickness of 1 ft above and 1 ft under the

geogrid necessary to eliminate the effect of theses boundaries on test results.

Modeling load transfer in soils reinforced with geogrids is complex and

involves several interaction mechanisms. However, a simplified load-transfer

model for extensible reinforcement was developed in order to rationally

extrapolate laboratory pull-out test results to full-scale reinforcement lengths.

Development of this model required the appropriate instrumentations to

measure the displacements at different locations along the geogrid. The data

analysis permitted the derivation of the soil-geogrid interaction parameters and

the in-soil stress-strain relationship of the reinforcement from displacement-rate

controlled pull-out tests. The confined creep characteristics were determined

from the interpretation of stepped load-controlled pull-out tests.

The methods currently used for the design of geosynthetic reinforced


199

walls and embankments are mostly derived from conventional slope stability

analysis modified to account for the reinforcement effect on the global stability

of the soil mass. These methods do not provide an estimation of the tension

forces mobilized at each reinforcement and, therefore, can not be used to

evaluate the local stability at each reinforcement level. These limitation raised

the need to development of a design approach that utilizes the interface

parameters and in-soil stress-strain characteristics derived from pull-out test

results . This approach considers the strain compatibility between the

reinforcement and the soil under working loads and allows the evaluation of the

effect of soil dilatancy and reinforcement extensibility on the generated tension

forces at each reinforcement level. The predicted tension forces and failure

angles are found to be consistent with the observations on reduced scale lab

models, full-scale model walls, and embankments with different types of

extensible reinforcements. The comparison between predicted and measured

tension forces in these walls demonstrates that the two assumptions,

commonly used to predict the state of stresses in the reinforcement (i.e. Ka and

Ko) are too restrictive to properly simulate the observed variation of tension

forces with depth for different extensible reinforcements.


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204

Juran, I., and Christopher, B., (1989), "Laboratory Model Study of Geosynthetic Reinforced Soil Retaining Walls," Journal of Geotech. Engineering., ASCE, Vol. 115, No. 7, pp. 905-926.

Juran, I., Ider, H., and Farrag, K., (1990), "Strain Compatibility Analysis for Geosynthetic Reinforced Soil Walls," Journal of Geotech. Engineering, ASCE, Vol. 116, No. 2, pp. 312-329.

Knochenmus, G., (1987), "Review of the Available Pull-out Testing Facility," Report submitted to Louisiana Transport. Research Center, Research Project 87-1G7.

Knochenmus, G., (1989), "Stress-Strain Behavior of Non-woven Geotextiles under Confined Conditions," Thesis presented to Louisiana State University for M.Sc. Degree, 194 p.

Koerner, B., (1986), "Direct Shear/Pull-out Tests on Geogrids," Report No. 1, Dept, of Civil Eng., Drexel Univ., Philadelphia, PA.

Leshchinsky, D., and Reinschmidt, A.J., (1985), "Stability of Membrane Reinforced Slopes," Journal of Geotech. Engineering, ASCE, Vol. 111, No. 11, pp. 1285-1300.

Leshchinsky, D., Volk, J.C., and Reinschmidt, A.J., (1986), "Stability of Geotextile-Reinforced Earth Railroad Embankments," Geotextiles and Geomembranes, 3, pp. 105-128.

Leshchinsky, D., and Perry, E.B., (1987), "A Design Procedure for Geotextile-Reinforced Walls," Geosynthetics Conference, New Orleans, pp. 95-107.

Leshchinsky, D., and Field, D.A., (1987), "In-Soil Load Elongation, Tensile Strength and Interface Friction of Non-Woven Geotextiles," Geosynthetics Conference, New Orleans, Vol. 1, pp. 238-249.

Leshchinsky, D., and Perry, E.B., (1989), " On the Design of Geosynthetic Reinforced Walls," Geotextile and Geomembranes, 8, pp. 311-323.

Leshchinsky, D., and Poedeker, R., (1989), "Geosynthetic Reinforced Soil Structures," Journal of Geotech. Engineering, ASCE, Vol. 115, No. 10, pp. 1459-1478.


205

Martin, J.P., Koerner, R.M., and Whitty, J.E., (1984), "Experimental Friction Evaluation of Slippage Between Geomembranes, Geotextiies and Soils," Proc. Int. Conf. on Geomembranes, Denver, pp. 191-196.

McGown, A., Andrawes, K.Z., and Kabir, M.H., (1982), "Load-Extension Testing of Geotextiles Confined In-Soil," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 793-798.

Milligan, V., and La Rochelle, P., (1984), "Design Methods for Embankments over weak soils," Polymer Grid Reinforcement, London, pp. 95-102.

Mitchell, J.K., Campanella, R.G., and Singh, A., (1968), "Soil Creep as a Rate Process," Journal of the Soil Mech. and Foundation Division, ASCE, pp. 231-253.

Mitchell J.K., (1969), "Shearing Resistance of Soils as a Rate Process," Journal of the Soil Mechanics and Foundation Division, ASCE, pp. 29-61.

Mitchell, J.K., and Villet, W.C., (1987), "Reinforcement of Earth Slopes and Embankments," Transport. Research Board, NCHRP Report No. 290.

Miyamori, T., Iwai, S., Makiuchi, K., (1986), "Frictional Characteristics of Non-Woven Fabrics," 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 3, pp. 701-705.

Murayama, S. and Shibata, T., (1961) "Rheological Properties of Clays," Proc. 5th. Int. Con. on Soil Mech. and Foundations, Paris, pp. 269-273.

Murray, R.T., (1982), "Fabric Reinforcement of Embankments and Cuttings," 2nd. Internat. Conference on Geotextiles, Las Vegas, V. 3, pp. 707-713.

Myles, B., (1982), "Assessment of Soil Fabric Friction by Means of Shear," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 787-791.

Myles, B., (1987), "A Review of Existing Tension Testing Methods," Geotextiles Testing and the Design Engineer, ASTM STP 952, pp. 57-68.

Palmeira, E., and Milligan, G., (1989), "Large Scale Direct Shear Tests on Reinforced Soil," Soils and Foundations, Japanese Society of Soil Mech. and Found. Eng., Vol. 29, No. 1, pp. 18-30.


206

Richards, E.A., and Scott, J.D., (1985), "Soil Geotextile Friction Properties," 2nd. Canadian Symp. on Geotextile & Geomembrane, pp. 13-24.

Richards, E.A., and Scott, J.D., (1986), "Stress-Strain Properties of Geotextiles," 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 3, pp. 873-878.

Rowe, R.K., (1984), "Reinforced Embankment, Analysis and Design," Journal of Geotech. Engineering, ASCE, Vol. 110, No. 2, pp. 231-246.

Rowe, R.K., Ho, S.K., and Fisher, D., (1985), "Determination of Soil- Geotextile Interface Strength Properties," 2nd. Canadian Symposium on Geotextiles, pp. 25-34.

Rowe, R.K., and Soderman, K.L., (1985), "An Approximate Method for Estimating the Stability of Geotextile-Reinforced Embankments'" Can. Geotech. Journal, Vol. 22, pp. 392-398.

Rowe, R.K., and Ho, S.K., (1986), "Determination of Geotextile Stress- Strain Characteristics using a Wide Strip Tests", 3rd. Internat. Conf. on Geotextiles, Vienna, pp. 885-890

Saxena, S.K., and Budiman, J.S., (1985), "Interface Response of Geotextiles," Proc. of the 11th. Internat. Conference on Soil Mechanics and Foundation Engineering, San Francisco, Vol. 3, pp. 1801-1804.

Shen, C.K., Romsted, K. and Hermann, L.R. (1976),"Integrated Study of Reinforced Earth, il-Behavior and Design," Journal of the Geotech. Div., ASCE, 102-GT6, pp. 577-590.

Siel, B.D., Tzong, W.H., and Chou, N.N., (1987), "In-Soil Stress-Strain Behavior of Geotextile," Geosynthetics Conference, New Orleans, Vol.1, pp. 260-265.

Singh, A. and Mitchell, J.K., (1968), "General Stress-strain-Time Functions for Soils", Journ. Soil. Mech. and Found. Division, ASCE, SM1, Jan., pp. 21-46.

Singh, A. and Mitchell, J.K., (1969), "Creep Potential and Creep Rupture of Soils," Proc. 7th. ICSMFE, Mexico, vol. 1.


207

Shrestha, S.C., and Bell, J.R., (1982-a), "A Wide Strip Tensile Test of Geotextiles," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 739-744.

Shrestha, S.C., and Bell, J.R., (1982-b), "Creep Behavior of Geotextile Under Sustained Loads," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 769-774.

Schlosser, F., and Elias, V., (1978), "Friction in Reinf. Earth," ASCE National Convention, Pittsburgh, PA., pp. 735-764.

Schlosser, F., and Guilloux, A., (1979) "Friction Between Soil and Strips in Reinforced Earth Structures," Intern. Conference of the Reinforcement of Soils, Paris,France, Vol. 1.

Schlosser, F., Jacobsen, H.M., and Juran, I., (1983),"SoilReinforcement," 8th. European Conference on Soil Mechanics and Foundation Engineering, Helsinki.

Schmertmann, G.R., Chouery, C., Johnson, R., and Bonaparte, R.,(1987) "Design Charts for Geogrid Reinforced Soil Slopes," Geosyunthetic Conference, New Orleans, pp. 108-120.

Solomone, W.G., Boutrup, E., Holtz, R.D., Kovacs, W.D., and Sulton, C.D., (1980), "Fabric Reinforcement Designed Against Pull-out," The Use of Geotextiles for Soil Improvement, ASCE, Portland, Oregon, pp. 80- 177.

Stewart, J.R., Williamson, and Mohnen, (1977), "Guidelines for use of Fabrics in Construction and Maintenance of Low-Volume Roads", USDA Forest Service, Portland, Oregon.

Tzong, W.H., and Cheng-Kuang, S., (1987), "Soil-Geotextile Interaction Mechanism in Pullout Test," Geosynthetics Conference, New Orleans, Vol. 1, pp. 250-259.

Vyalov, S.S., (1986), "Rheological Fundamentals of Soil Mechanics," Developments in Geotechnical Engineering 36, Elsevier.

Williams, N.D., and Houlihan, M.F., (1987), "Evaluation of Interface Friction Properties Between Geosynthetics and Soils," Geosynthetics Conference, New Orleans, Vol. 2, pp. 616-627.


APPENDIX A

EQUIPMENT DETAILS & SPECIFICATIONS.

208


209

- o

Figure A -1. Top View of the Pull-out Box


210

H

H

Figure A-2. View of the Pull-out Box.


211

Figure A-3. Detail of the Loading Frame.


212

Figure A-4. Detail of the Loading Frame.


K $ \

-4 -

Figure A-5. Detail of Load Cell Connections.


nrjf do

Ola

214

4 - & d . a-.y.4 -

GD— IWCET

Figure A-6. Detail of Hopper.


Si F £ £ =* 115 * s

V1 lii i .

5 !1

Figure A-7. Specifications of the LVDT's.


216

STRAIN GAGE LOAD CELL MODEL 3187 - 20K S/N 751___________ CALIBRATION DATE 3/19/90

SPECIFICATIONS:RATED CAPACITY-TENSION & COMPRESSION... ■ 20,000__Lbs.____________MAX. LOAD (without zero shift)........ 50% overload (150% of rated

capacitySIGNAL SENSOR............................ 4 arm bonded strain gage bridgeBRIDGE RESISTANCE........................350 ohms nominalMAX. BRIDGE EXCITATION..................20 volts DC or AC RMSCOMPENSATED TEMP. RANGE.................70° F. to 170° F.USEABLE TEMP. RANGE..................... -65° F. to +200° F.

•EFFECT OF TEMP. ON ZERO................. ±0.002% of rated capacity/°F.•EFFECT OF TEMP. ON OUTPUT............... ±0.002% of reading/°F.NONLINEARITY............................. ±0. % of rated capacity 'OUTPUT.................................... COMP. - 2.998 mV/V at rated cap.

TENS. 4 3.003 mV/V at rated cap.•Within compensated temperature range

ELECTRICAL CONNECTIONS:RECEPTACLE: PT02E-10-GP or MS-3102 A-14S-5P PIGMATING CONNECTOR: PT06W-10-6S or MS-3106 E-14S-5S TAILS

PINS FUNCTION RESISTANCERED A (+) and (-) D BLK Excitation 358.3 ohmsGRN B (+) and (-) C WHT Signal ■ 350.7 ohms\ '

CALIBRATION:A precision wire wound resistor, when shunted across one leg of the strain gage bridoe, produces an electrical signal equivalent to an applied load. This shunt calibration is valid only when used with high input impedance indicators. The equivalent values below were determined by factory calibration.

LOAD VALUE ACROSS PINS RESISTOR VALUE14,539 Lbs, compression GRN B and D BLK 40 K ohms14,526 Lbs, tension GRN B and A RED 40 K ohms

Figure A-8. Specifications of the Load Cell.


217

Kj’Mearth preiture) ’/•"•(concreie ttrrts)

—j~mn i r'T -rTunjdueer Houiini

Flexible Conduit Cable

i ..limits

Minrmurn*l2''-Miximum-20’*-

Geokon offers the following four basic types o f transducers:

Pneum atic (Models 2510 and 25201. T h is transducer is a Petur M odel P -100 and is connected to the readout lo ca tion via tw in pneumatic tubes (M odel T -102). I t Is read ou t using the Petur M odel C-102 Readout Box.

S em iconductor S tra in Gage — C onventiona l style pressure trans* ducer com patible w ith many ex is ting dataloggers where 5 vo lt DC or AC exc ita tion is provided. The signal o u tp u t is large enough to be scanned d ire c tly w ith o u t fu rth e r am p lifica tion . They may also be read rem ote ly using the G eokon M odel RB- 101 Readout Box.

Resistance Strain Gage - These itansducers are com patib le w ith o the r resistance strain gage systems such as load cells, borehole de fo rm a tion gages etc. Ih c y may be read out using convention* a ls tra in ind ica tor o r readout boxes(V ishay P350A, Vishay P3500 e tc.).

V ib ra tin g Wire (Model 4500H ). This transducer is com patib le w ith the rest o f the G eokon line o f vibrating wire instrum ents and is recommended fo r use w ith the Model GK*4 Datalogger or GK*401 Readout Box. I t incorporates all o f the advantages o f the v ib ra ting wire system o f measurements.

S P E C IF IC A T IO N SModel Number EP Cell

CS Cell2510

25203500

36003650

36604800E

4800CTransducer type — Pneumatic

Semiconductor strain gage

Resistance strain gage

Vibrating wire

Typical ranges available psi15 to 3000*

25,100,500 25(100)500 50,100,500 1000,3000,5000 10003000,5000 1000,3000,5000

Over-range capacity , % F.S. 3000 psi max. 200 200 . 150Accuracy '% F.S. 0 .4 * 1 0.5 0.25Resolution % F.S.

*X}“O

infinite infinite 0.1Thermal effect on aero % F .S./°F aero <0 .05 <0.02 <0.05Excitation voltage V — max.6v

(DC or AC)✓m^x.lOv

(fpCJor AC)5v sq.wave

Signal output mv/v — 20 3 12UO-2000Hafrequency

Bridge resistance ohms - 150in 1 15out 350 coil 150Transducer housing dia. in. 1.5 1.625 2.25 1Transducer housing length in. 6 6 6 6Weight (less cable lbs. 5 7 5 7 5 7 5 7Connecting leads — Polyethylene

twin tubes4-cond.shielded

4-cond. • shielded

2-cond.shielded

Readout box - Petur C-102 Geokon RB-101 Vishay P350A Geokon GK-401"Depends on pressure gage in readout box.

Figure A-9. Specifications of the Pressure Gauges.


APPENDIX B

COMPACTION CONTROL AND DENSITY CALIBRATION.

218


Figure B-1. Points od Density Measurements.

Reproduced with permission of the copyright owner. Further reproduction prohibited without permission

220

I M II II II II II II II II II II II II II II II II II II II II It II II II II II II II II II II II II || II II II II II II It II || II II II II II II II II II It II II II II II II II II II || ||

pt *' blows " layer 2 " layer 3 up " layer 3 down " layer 4

II •» II II

0 II 99. 1 ( 0 ) it 101 .9 (+0.8) II 102.6 (+1.0) II 103.840 II 102 . 3 (+0.6) II 101 .9 (-0.7) II 104.4 (+0.8) II 104.7

110 '* 103.8 (+0.3) II 105.3 (+1.0) II 106.2 (+1.1) II 106.7200

II104. 1 (-0.1) It 108.0 (+2.9) • 1

II108.2 (+1.2)

II108.4

II II II II

0 II 99.3 (+0.2) II 102.2 (+1.1) • 1 102.7 (+1.1) II 103.540 II 102.4 (+0.7) II 102.6 ( 0 ) II 104.3 (+0.7) II 105.9

1 10 II 103.2 (-0.3) II 103.4 (-0.9) It 104.0 (-0.9) II 107.5200 II

II104.7 (+0 . 5) It 104.3 (-0.8) II

•1105.7 (-1.3) II

II10B.9

II II • 1 II

0 " 98.4 (-0.7) II 101 .4 (+0.3) II 101 .6 (+0.5) II 103.240 II 100.8 (-0.9) II 101 .7 (-0.9) It 102.2 (-0.4) 105.2

110 II 103.4 (-0 . 1) M 104.3 ( 0 ) II 104.6 (+0.3) II 107.0200 II 104.3 (+0.1) II

II103.1 (-2.0) ■ 1

•1105.8 (+0.7) II

II108.0

0IIII 98.7 (-0.4)

ItII 99.9 (-1.2)

It

•I 100.6 (-1.0)II

II 102.940 II 100.7 ('-1.0) II 102.5 (-0.1) •1 102.8 (-0.8) II 104.7

110 II 103. 1 (-0.4) II 103; 9 (-0.4) II 105.7 (+0.6) II 106 . 1200 •I

II103.4 (-0.8) II

II104.0 (-1.1) II

• 1106.3 (-0.7) It

II107.4

0II

II 100.6 (+1.5)II

II 101.3 (+0.2)II

102.2 (+0.6)IIII 103.2

40 II 102.6 (+0.9) II 104.9 (+2.3) II 104.4 (+0.8) II 105.0110 II 103.B (+0.3) II 105.2 (+0.9) II 105.5 (+0.4) II 107.5200 .. 103.6 (-0.6) II

II106.5 (+1.4) II

II107.9 (+0.9) II 109.1

0IIII 98.5 (-0.6) II 99.8 (-1.3)

IIII 99.6 (-2.0)

liII 103.0

40 II 101.8 (+0.1) II 102.2 (-0.4) II 103.5 (-0.1) II 103.9110 II 103.8 (+0.3) II 103.5 (-0.8) II 104.5 (-0.6) II 104.6200 II 104.8 (+0.6) II 104.7 (-0.4) II 107.8 (+0.8) It 109.9

II II II IIIt II II M

(+0.6) (+0.4)

( 0 > (+0.1)

(+0.6)(+0.6)

(+1.4)

Table B-1. Density Measurements at Each Layer.(Variable Falling Height).


221

Table B-1 (Continued)..

pt blows layer 2 "II II II II II It II II •< II II II M It II II II II || II II || || || ||

7 0 /40 /

110 /200 /

8 0 /40 /

110 /200 /

9 0 /40 /

110 /200 /

layer 3 up layer 3 dow nI II M II II II II II II II II II II It II II It II II II II II II II II II II It II I

It I

98.9 (—2. E > " 104.6 (-t-3.0) ’104.6 (+2.0) " 105.8 (+2.2) '106.3 (+2.0) " 107.2 (+2.1) •106.4 (+1.3) " 107.4 (+0.4) '

102.8 (+1.7) " 104.1 (+2.5)105.1 (+2.5) " 105.5 (+1.9)106.9 (+2.6) " 107.4 (+2.3)107.8 (+2.7) " 108.1 (+1.1)

102.5 (+1.4) " 104.2 (+2.6)103.9 (+1.3) " 105.6 (+2.0)104.5 (+0.2) " 107.5 (+2.4)106.4 (+1.3) " 107.9 (+0.9)

layer 4t II II II II It II It II ti ii il it

104.6 (+0.9)106.1 (+0.9) 107.8 (+0.9)108.1 (-0.4)

104.7 (+1.0) 105.2 ( 0 )107.4 (+0.5) 10B.5 ( 0 )

104.3 (+0.6) 105.7 (+0.5) 107.5 (+0.6)108.3 (-0.2)


222

II II II II II II tl ll ll ll 1 it ll II II ll ll l II ll ll ll it ll ll II ll ll ll ll ll l II II II II II M ... . •• 1 ll ll ll ll tl it || ll ll *1 II II II I ll ll ll ll ll ll

pt b lows layer 2 layer 3 up 1 aver 3 down layer 4•1 II II it H ll ll ll ll ll ll it ll ll ii it i ll II II II ll tl tl ll ll ll it ii ii n II ll ll It ll II *i II II ll ll II 1 ll II ll ll ll ll li II ll ll ll ll ll .......1 0 101 . 1 (-0.8) 101.9 (-0.9) 1 04. 1 (+0.8) 104.6 (+0.5)

40 103.8 ( 0 > 104.3 (+0 . 2) 105.7 (+0.7) 105.9 (+0. 1 )110 105.8 (+0.5) 106. 1 (+0.1) 106.6 (-0.3) 107.0 (-0.4)200 106.0 (-0.3) 105.2 (-1.6) 108.0 ( 0 ) 108.1 (-0.9)

a 0 101 .7 (-0.2) 103.1 (+0.3) 102.9 (-0.4) 103.8 (-0.3)40 104.7 (+0.9) 104.6 (+0.5) 105.3 (+0.3) 105.5 (-0.3)

110 105.5 (+0.2) 106.3 (+0.3) 107.3 (+0.4) 107.3 (-0. 1 )200 105.2 (-0 . 9) 107.4 (+0 . 6) 107 . 6 (-0.4) 108.7 (-0.3)

3 0 102.5 (+0.6) 103.8 (+1.0) 103.5 (+0.2) 103.9 (-0.2)40 104. 1 (+0 . 3) 103.9 (-0.2) 104.5 (-0.5) 105.6 (-0.2)

110 104.9 (-0.4) 105.3 (-0.7) 106.4 (-0.5) 106.9 (-0.5)200 106.0 (-0.3) 105.8 (-1.0) 108.4 (+0.4) 107 . 2 (-1.8)

4 0 102.4 (+0.5) 102.9 (+0.1) 103.7 (+0.4) 104.7 (+0.6)40 103. 1 (-0.7) 105.0 (+0.9) 104.9 (-0.1) 106.4 (+0.6)110 105.3 ( 0 ) 105.5 (-0.5) 106. B (-0.1) 107. 1 (-0.3)200 108.2 (+1.9) 106.7 (-0.1) 109 . 1 (+1.1) 109.6 (+0.6)

5 0 102. 1 (+0 . 2) 102 .2 (-0.6) 102.6 (-0.7) 104.2 (+0.1)40 103.5 (-0 . 3) 103.5 (-0.6) 104.6 (-0.4) 106.1 (+0.3)110 106.0 (+0.7) 106.8 (+0.8) 107.5 (+0.6) 107. 1 (-0.3)20 107.2 (+0.9) 108.0 (+1.2) 107.7 (-0.3) 1 10.0 (+1.0)

6 0 101 .5 (-0.4) 102 . 7 (-0.1) 103.2 (-0 . 1 ) 103.5 (-0.6)40 103.4 (-0.4) 103.2 (-0.9) 105.2 (+0.2) 105.2 (-0.6)1 10 104. 1 (-1.2) 106.7 (+0.7) 106.8 (-0.1) 107.5 (+0.1)200 105.2 (-1.1) 107.5 (+0.7) 107.1 (-0.9) 108.3 (-0.7)

Table B-2. Density Measurements at Each Layer. (Constant Falling Height).


223

Table B-2 (Continued).

II II II ll ll ll ll II ll ll ll •I n it it ii ii ii H ii ii ii M M ii ll ll ll ll ii ii ii II It II II II ll •• 1ll II II II II ll II II ll ll ll ll ll ll It tl tl ll ll ll it •1 ll II tl ll ll

pt blows "II

layer E layer 3 up ' layer 3 down layer 4II II II ll ll ll ll ll ll M It ll ll ll ll It ll tl ll ll ll ll ll ll II ll ll ll ll ii ii it II II 11 II II II II 1ll ll H ll ll ll HIi ll ll ll it t« it •I ll ll ll It ii ll II it ll It II ll

7 0 / 103.9 (+1.1) ' 103.8 (+0.5) 103.9 (-0.5)40 / 103.9 (-o .a > 1 105.7 (+0.5) 105.5 (-0.3)110 " / 105.3 (-0.7) ' 107.5 (+0.6) 108.3 (+0.9)EOO

tl/ 108.3 (+1.5) * 109.1 (+1.1) 110.0 (+1.0)

8II

0 / 104. 1 (+1.3) ‘ 103.9 (+0.6) 104.7 (+0.6)AO “ / 104.3 (+0.5) ' 106.4 (+1.4) 106.7 (+0.9)

110 " / 107.5 (+1.5) 1 108.3 (+1.4) 107.5 (+0.1)200

ii

/ 107.6 (+0.8) * 107.8 (-0.5) 110.4 (+1.4)

9ii

0 " / 103.7 (+0.9) 1 104.5 (+1.5) 103.6 (+0.5)AO / 104.4 (+0.3) ' 106.0 (+1.0) 105. 1 (-0.7)110 " / 108.4 (+5.4) ' 108.5 (+1.3) 107.6 (+0.5)200

H

n

106.8 ( 0 ) • 107.3 (-0.7) 109. 1 (+0.1)


APPENDIX C

PULL-OUT TEST RESULTS.

224


75

TEST#1

60

45

30

15

Un c o n f i n e d E x t e n s i o n0

0 too 16060 tooDisplacement (mm)


75 226

60

Test*2

45

■oMo►J

30

15


Dl«pUc»ment (mm)

(a )

TEST #2

160

008

tooClMI

mm/secSlope mm/min)

IM 1600 •000 •M0 •000Time (oeo)


Dis

plac

emen

t (m

m)

Load

(KN/

m)

227

75

TEST#3

60

45

30

15

Uncon f i n ed Ex t e ns i on

0 60 100 160Displacement (mm)

(a)

WOO

TEST #3

160

100

60

Slope= 0.05 mm/sec (3. mm/min

90001600 90000 600 1000

Time (sec)


Dis

plac

emen

t (m

m)

Load

(K

N/m

)

228

TEST-C1

Confin ing pressure3 7. psi

o M too 160 •00Front Displacement (nun)

(a)

Test-C l

160

Front D is p l .

100

Back D i s p l .

0 •00 400 •00 •00 1000Time (sec)


TEST-C2

Confining pressure- 7 . psi

I N

Front Displacement (mm)IM

(a )

TEST-C2

ISOFront D is p l.

aeavu«j

Back D is p l.too

aJ)Q

•00000•00 400Time [sec]

(b )


23090'

TEST-C3

30

psiConfining pressure

190 90990 1900Front Displacement [mm]

(a )

TEST-C3

190

Front D isp l.i«w*

aVSvua

100

ans

Slope- 0 .18 mm/sec (1. cm/min)

1000Tima (tec)

(b )


TEST-C4

B 45

•o 30

.Confining pressure- 7 . psi

INIMFront Displacement (mm)

(a )

TEST-C4

IM

ront D isp l.

1I

IM -CJ

I

Slope- 0 .18 mm/sec (1. cm/min)

m mto* m I M

Time (eeo)

(b )


150

TEST-A1

120i

30

Confining pressure- 10. psi

0 20 40 60 BO 100Front Displacement (mm)

(a )

IM

TEST-A1

Front D isp l.

1aeue

Node

•M ■NMTime (aee)

(b )


232

150

TEST-A2190

60

Confining Pressure- 7. psi

100606040200Front Displacement [m m ]

(a )

100

TEST-A280

60dvQVo0&s

40

S lope- 0 .1 mm/sec

0 200 400 600 600 1000Time (eec)

(b )


90

•oao

75

60

TEST-A3

30

15

Confining pressure- 7 . psi

0200100 150500

Front Displacement (mm)(a )

200

TEST-A3Front D is p l.

150

0V0Vo0P<V)

Node (4)100

Node (5)o50

0 500 1000 1500 2000Time (aec)

(b )


90

TEST-A4

60

43

15

Confining pressure** 7. psio

0 50 100 150 200Front Displacement [mm]

(a)

200TEST-A4

150 Front Displ.

tiV0VV 0Aj2

100

Node (5)

D

50

1500 20001000‘ 5000Time (sec)

(b)


90 236

75

60

TEST-A5

Confining pressure* 7 . psi

50 100Front Displacement [mm]

150 200

(a )

200

TEST-A5

150Front D isp l.

Node (1)

Node (2)flVaVo0an

100

a

50

0 500 1000 1500 2000Time (sec)

(b)


90

TEST-A688

68

45

30

15

Confining pressure - 5. psi

0150 200100500

Front Displacement (mm)

(a )

150

TEST-A6 Front D is p l.

Node (1)

- (3)100

fl'a2

ao

50

0 500 1000 1500Front Displacement (mm)


90

TEST-A775

60

45

30

15

Confining pressure « 7 . psl

06o 2001501000

Displacement (mm)(a )

200 1TEST-A7

150 Front D is p l.

NodeNode

Node

50

Node (5]

’ 5000 1000 1500 2000Time (Sec)

(b )


TEST-AB

OU

^ \ sB

100Front Diaplacement (mm)

150 200

(a )

2 00

TEST-AB

160 Front D is p l. Node (1)

NodeNodeNode100

50

00 500 1000 1500 2000

Time (aec)

(b )


TEST-A9

S 30

Confining pressure - 7 . psi

50 150 8000 100Frost Displacement (mm)

(a )

200

TEST-A9Front D is p l. Node (1)150

Node

Node0 100

50Node (5j

0 500 1000 1500 2000Time (eec)

(b )


TEST-A1060 -

/■Nai•Oao

0 SO 40 00 00 100 140 100120 100

Fact Dlsplacaaant M(a )


90

TEST-B175

60

30

15

Soil Thickness ■ 40 cmo

20015050 1000F ro n t D isp lacem ent (m m )

(a )

200

TEST-B1

Front D is p l.150

NodeNode

Node (4)§ 100

0 1500500 1000 2000Time [Bee]

(b )


Dis

plac

emen

t (m

m)

90

75 -

B

T3ao■-]

15

TEST-B2

.60

45

30

S o il Thickness - 40 cm

50 100F ro n t D isplacem ent [m m ]

160 200

(a)

200

TEST-B2

Front D isp l.150

Node

Node

Node

60

00 boo 1000 1500 2000

Time (sec)


90

TEST-B375

60

45

30

15

Soil Thickness - 60 cmo

0 50 100 150 200F ro n t D isplacem ent (m m )

(a )

200

TEST-B3Front D isp l.

150NodeNodeNode

aClauo

8Q60

20001500'500 1000 Time (sec)

<b)


90

TEST-B475

60

45

30

15Soil Thickness ■ 20 cm

o60 200100 1600

F ro n t D isp lacem en t [m m ]

(a )


90

TEST-P175

60

55

30

15

Sleeve Length = .12 in0

1000 150 20050F ro n t D isp lacem ent [m m ]

(a )

200

TEST-P1Front D is p l.

150Node

Node

100oan..gn

50

100 ‘ 300 400 500 6000 200Time ( b o o )

(b )


/•NB

•OWOhJ

90

TEST-P275

60

30

15

00 50 100 150 200

247

F ro n t D isp lacem ent [m m ]

(a )

200

TEST-P2

Front D isp l.150

Q 100 Node (4)

100 200 300 Time [sec]

400 600 600

(b )


Dis

plac

emen

t (m

m)

90

B

75

60

65

•awo^ 30

15 .

TEST-P3

60 100 150 200F ro n t D isp lacem ent (m m )

(a)

200

TEST-P3

Front D is p l.150

Node (4)

100

50

00 100 200 300 400 500 600

Tim e [eec]

n . 0 0


90

TEST: L0AD-U175

60

•o 30

15

Unconfined Extensiono

400 500200 3000 100Time (m ln)

(a )

200

TEST: L0AD-U1

150

loO0 200 300 400 500Time (mizx)

(b)


Dis

plac

emen

t (m

m)

TJ<9

90

TEST: L0AD-U275

60

5

0t'

15

Unconfined Extensiono

o 100 200 300 600400

250

Time (m in)

(a )

200

TEST: L0AD-U2

150

100

50

01000 300 400 500200

Time (m in)


75

60

Unconfined Extension Test

Geogrid: Conwed

Test: C0N-U1

45

« 30o►j

15

to 40 eoDisplacement (mm)

BO 100

Unconfined Extension Test

Geogrld: Conwed

Test: C0N-U2

S 30

5 15

100806040200D isp lacem en t (m m )


Load

(K

N/m

)

252

Unconfined Extension Tests

Geogrid: ConMed60Test: Con-U3

0 20 40 60 80 100D isplacem ent (mm)


75

60

45

30

Test: CON-1

Geogrid: Conwed15

Confining Pressure - 7 . p s i

o 1008060400 20F ro n t D isplacem ent (mm)

100

Test: CON-1

80EE

n Nodec©E©U6 Node #1a09aB•ao2

Node #2

20 Node #3

Node #4 Node #5

250020001500 30000 1000500Time (sec)

Reproduced with permission of the copyright owner. Further reproduction prohibited without permission

254

6 0 '

.*>a

s30

Test: CON-2

Geogrid: Conwed

Confining Pressure - 7 . psi

10060 800 20 40Front D isplacem ent (mm)

100

Test: CON-2

80EE

a 60VE©©J3aJ2 40 Q

Node #0

Node #1

a•aoZ

Node #2

20 Node #3

Node #4 Node #5

300025002000150010005000Time (sec)


75

60

45

30

Test: CON-3

Geogrid: Conwed15

Confining Pressure ■ 7 . psi

01000 6020 8040

F ron t D isplacem ent (mm)

100

Test: C0N-3

60

B-B 60Node #0

m*a

io■g- 40s Node #1

Node #2

20Node #3

Node #4

9000250020001600500 10000lim e (sec)


126

105

B8"■s

■oflo■J

Test: C0N-P1

Geogrid: Conwed

C onfining Pressure - 20. psi

25 30200 5 10 15F ro st D isplacem ent (mm)

40

Test: C0N-P1

? 30

20Node #0

CL

•OS5 10

Node #1

Node #2

Nodei *4 -5

1000 15000 500Time (sec)


257

105

Test: C0N-P2

Geogrid: ConMed

Confining Pressure * 20 . psi

30o 5 15 20 2510Front D isplacem ent (mm)

Test: C0N-P2

Node *0so

■c£S 20ac-ua

oSc

lode #1

lode #2

1000BOOBOO4000 zooTim* (>*e)


120

105

Test: C0N-P3

Gegrid: Conned

Confining Pressure - 20. psi

300 5 10 15 20 25F ront D isplacem ent (m m )

40

Test: Con-P3

Node #oE 30

20

2 10 Node #1

Node #2

10008006004002000Tim e (se c )


APPENDIX D

FORMULATION OF SOIL-REINFORCEMENT

MODEL RELATIONSHIPS.

259


260(I) soil Yield Function:

The soil is assumed to follow Mohr-Coulomb yield criteria:

Tcv = h (Kxy) • a r

where and ay are the soil mobilized shear and normal stresses along the

failure surface, h(yxv) is a strain hardening (and softening) function relating the

current yield surface to the actual state of strain. As soil cohesion c = 0, the

function h(Kxy) = tan <pm , where 0 m is the mobilized soil friction along the

failure surface, then:

"xy = (Kxy)tfy

and the yield function (Figure D.1) is:

F(a.Kxy) = — - h(Kxy) = 0 (D.1)

For dense dilating sand, the strain hardening (strain softening) function h(Kxy)

is assumed parabolic; i.e.:

c • KxV (Kx.v - a)h(Kxy) = tan 0 m = ........................... (D.2)

(Kxy + b)2

The constants a, b and c are determined from the initial conditions as follows:

(i) the initial tangent of h(yxy) = G/ay ; i.e.:

d[h(Kxy)] G------------- = — (at yxv = 0)

b K x y O -y


261

(Kxy + b)2 (2c• Kxy - c.a) - 2 c.y^ (Kxv ' a) (K*y + b) G

(K*y + b)4 <rv

- c.a G = — (D.3-a)

(ii) at peak shear strain, h ^ ) = tan 0 P

d[h(Kxv)land ------------ = 0 (for Kxy = KP)

dKxy

i.e., (kp + b) (2.c.kp) = 2 .c .kp (kp - a)

a.bKP = ---------

2b + a

- a2, bsubstituting, h(Kp) = ----------------- = tan 0P (D.3-b)

4b (a + b)

(iii) at critical state, l im ^ h(Kxy) = tan 0 CV

c-K2xv - a.c.Kxvi.e., lim ^. h(Kxy) = lim^.. [ ---------------------------------] = c

y2 xy + 2.b.Kxy + b2

c = tan 0 CV (D.3-c)

solving equations (D.3-a), (D.3-b) and (D.3-c) we get:

b = 2 (ay/G). tan 0 P . [1. + 1.-(tan Q J tan 0 P) ]

a = -4 (Oy/G). tan20p . [1. + 1 .-(tan Q J tan 0P) ]2/ tan 0 CV


262(II) Soil-Reinforcement Interaction:

From the deformation in reinforcement during shearing, Figure (D.2), we get:

a + Ax a/h + Ax/htan A, = -------------- =--------------------

h + Ay 1 + Ay/h

Then,tan + dKxy

tan A, - ........................... (D.4)1 + dey

From Mohr's circle of strain (Figure D.3), we get

deR = dey + x = deR + L sin A

Where deR is the incremental strain in the reinforcement. From Figure (D.3) we

get : L/ sin y = dey/ sin u

and sin y = cos (A + u) , substituting:

L = cos M + </). dey / sin {u)

then,cos (A+u). sin A

deR = dey + ( ------------------------- ) deysin u

since dey = tan u. dy^ , we get:

cos (A +t/). sin AdeR = tan u ( 1. + -------------------------- ) dy^ .......................... (D.5)

sin u

(III) Equilibrium Equations:

For a potential failure surface a, Figure (D.4) shows the forces acting on the

active zone.


263The horizontal equilibrium for any slice J (j = 1,2...N), neglecting

the forces between slices, is:

Tj .cos P, = Rj .sin (a - 0 )

where Tj is the tension force of the reinforcement at level (j) = E.t.b P is

the reinforcement inclination with the horizontal, and cos P = sin Mj + a), and

Rj is the soil reaction along the failure surface.

.E.t.b. sin (At + a) = Rj. sin (a-0)

then,sinMj +a)

ZNRj = E.t.b. I N fpj . ................... (D.6)sin(a-#j)

The weight of the soil mass in the active zone (Figure D.4) is:

W = 'A p . H2 [ (1/tan a) - tan J ] . Sh (D.7)

where H is the wall height, J is the inclination of the wall facing with vertical,

and Sh is the horizontal spacing of the reinforcement. The vertical equilibrium

in the active zone is:

I N Rj cos (or-0j) - I N Tj sin P « W (D.8)

Substituting equation (D.6) into equation (D.8), we get:

sinMj + a)E.t.b.{ I N fpj.---------------- - I N cosMj + o) } = W

tan(a-0j)

Substituting equation (D.7) and rearranging the equation, we get the non-

dimensional relationship:


264E.t.b 2 [ I N sinMj + a) cot(a-0) -ZN € cosMj + a)]

H /[--------------] ,= . . . (D.9)p.Sh.Sv N[(1/tan a) - tan J]

The most critical solution in terms of the critical height H under a specific shear

strain yxy can then be obtained by minimizing H in equation (D.9) with respect

to the failure angle a . These formulations are implemented in a computer

program to obtain the critical height and the tension force distribution at each

reinforcement level for incremental shear strains.


265

tan ^

tT _ ton <f>{U

4T

W

Figure D.1 Elasto-plastic Strain Hardening function of the Soil.

Figure D.2 Deformation of the Reinforcement during Shearing.


DIRECTION OF R E IN FO R C E M E N T

dc

L U L L

ZERO EXTENS LINE (Horizon!

Figure D.3 Mohr's Circle of Strain Increments in Soil.

Figure D.4 Failure Mechanism in Reinforced Soil Walls.


VITA

Mr. Khalid A. Farrag was born in Alexandria, Egypt on Aug. 15, 1955.

He got his B.S. degree in Civil Engineering from Cairo University on 1977. He

worked as a Civil Engineer in Egypt and Saudi Arabia from 1978 till 1984. He

Got his M.S. degree in Civii Engineering from Louisiana State University on Dec.

1986, and his Ph.D. degree in Civil Engineering on Dec. 1990.

267


DOCTORAL EXAM INATION AND DISSERTATION REPORT

Candidate: K h a lid A. Farrag

Major Field: C iv i l E ng ineering

Title of Dissertation: in te r a c t io n P ro p e rtie s o f Geogrids in R e in fo rced S o il W a lls ,

T e s tin g and A n a ly s is .

Approved:

M ajor Professor and Chairman

Dean of the Graduate School

EX A M IN IN G C O M M ITTEE:

Date of Examination:

Nov. 19, 1990


Documents

Interaction Properties of Geogrids in Reinforced Soil