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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
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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
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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
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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
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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
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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
6.1 INTRODUCTION. 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.
7.1 INTRODUCTION. 159
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
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(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
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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
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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
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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
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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
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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
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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'.
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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.
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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
7.13 Measured and Predicted Tension Forces in Wall No. 2. 182
7.14 Geometry and Soil Properties of Model Walls. 185
7.15 Measured and Predicted Tension Forces in Wall No. 2. 186
7.16 Measured and Predicted Tension Forces in Wall No. 3. 187
7.17 Measured and Predicted Tension Forces in Wall No. 7. 188
7.18 State of Stress in the Reinforcement with Depth. 189
7.19 Measured and Predicted Tension Forces in Embankment 1. 191
7.20 Measured and Predicted Tension Forces in Embankment 2. 192
7.21 Measured and Predicted Tension Forces in Embankment 3. 193
7.22 Measured and Predicted Tension Forces in Embankment 4. 194
7.23 Measured and Predicted Failure Surface in Embankment 1. 195
xiv
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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.
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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.
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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
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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.
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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,
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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
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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
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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.
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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
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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.
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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.
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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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
■a- O
7T7 "OT
Figure 2.4 Confined Extension Testing Apparatus. (After McGown et al., 1982)
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14
Soli
- - Bottom Clomp
— ~ ~ Boss Plolo
(T) Pon Picitura © Cell Praaaura (3) Back Pinturi
Figure 2.5 Confined Extension Testing Apparatus.(After Knochenmus, 1989}
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
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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
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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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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]
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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].
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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.
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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).
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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)
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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].
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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].
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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.
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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.
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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.
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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
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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.
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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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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
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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.
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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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
1) Pull-out Box2) Loading Frame3) Elevated Sand Container.
Figure 3.1 Side View of the Pull-out Box.
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1) Loading Frame2) Clamping Plates3) Load Cell4) Geogrid Reinforcement.
Figure 3 .2 Front View of the Puli-out Box and Loading Frame.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
Figure 3.5 View of the Large Direct Shear Box.
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42
Air bag
Geogrid
Compacted sand
5 ‘ - O'
Figure 3.6 Cross-Section of the Direct Shear Box.
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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.
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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.
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4 5
1) Hydraulic Ram2) Loading Frame3) Load Cell.
Figure 3.7 View of the Loading Frame.
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A) Control BoxB) Fluid ReservoirC) Hydraulic RamD) Servo ValveE) Low Pressure Gauge.
Figure 3.8 Details of the Hydraulic loading System.
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A) Elevated HopperB) Vacuum SystemC) Hopper Outlet.
Figure 3 .9 View of the Sand Handling Facility.
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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,
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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,
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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.
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A) Stabilized Power SupplyB) Screw Terminal BoardC) Load Cell Conditioner.
Figure 3.10 Data Acquisition Supporting Instruments.
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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.
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^ocfuced
uCJQ.
CD
CD
in
h-o
Pull-Out Box
Figure 3 .14 Schematic View of the Data Acquisition System.
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Figure 3.15 View of the data Control System.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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m
ah*•*■* ait
ill!f i l la:cs
Figure 4.1 Grain Size Distribution of the Blasting Sand.
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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.
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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.
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Figure 4 .4 Soil Density Measurements in the Pull-out Box.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Unconfined Extension
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
91
(o)
(b)r
<c)
r
T T “ wr
Figure 4.20 Concept of Strain Superposition in Stepped Loading Tests.(After Vyalov, 1986).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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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Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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),
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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).
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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.
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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)
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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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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
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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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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.
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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
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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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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-
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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Load
(K
N/m
)128
120
90
60
30
10 cm above Sc 10 cm under the geogrid
10 cm above Sc 30 cm under the geogrid
30 cm above Sc 30 cm under the geogrid
40 cm above Sc 40 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.
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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
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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
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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
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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
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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
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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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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)
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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)
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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
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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'.
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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.
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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.
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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'.
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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'.
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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'.
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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.
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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'
Unconfined Extension
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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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'.
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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\
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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'.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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\
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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.
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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
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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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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'
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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'.
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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)
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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'.
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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
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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
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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
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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-
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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
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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
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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
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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
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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 .
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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).
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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.
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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)
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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
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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
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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
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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.
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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.
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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.
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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).
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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.
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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).
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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
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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.
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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.
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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.
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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)
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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.
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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)
Figure 7.15 Measured and Predicted Maximum Tension Forces in Wall No. 2.
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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)
Figure 7.16 Measured and Predicted Maximum Tension Forces in Wall No. 3.
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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)
Figure 7.17 Measured and Predicted Maximum Tension Forces in Wall No. 7.
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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.
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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).
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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.
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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)
Figure 7.20 Measured & Predicted Tension Forces in Embankment 2.
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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)
Figure 7.21 Measured & Predicted Tension Forces in Embankment 3.
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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)
Figure 7.22 Measured & Predicted Tension Forces in Embankment 4.
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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,
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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
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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,
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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
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
REFERENCES
Al-Hussaini, M.M ., (1977), "Field Experiment of Reinforced Earth Wall," Internat. Conference on the Use of Fabrics in Geotechnics, Paris, Vol. 1, pp. 119-121.
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Brakel, J., Coppers, M., Maagdenberg, A.C., and Risseuw, P., (1982), "Stability of Slopes Constructed with Polymer Reinforced Fabric, Test Section of Aimer, Holland," 2nd. Internat. Conference on Geotextiles, Las Vegas , V. 3, pp. 727-732.
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Christopher, B.R., (1976),"Tensar SS2 Geogrid Evaluation", Evaluation Report to Tensar Cooperation.
200
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201
Christopher, B.R., Holtz, R.D., and Bell, W.D., (1986), "New Tests for Determining the In-Soil Stress-Strain Properties of Geotextiles" 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 3, pp. 683-688.
Christopher, B.R., Safder, A.G., Giroud, J.P., Juran, I., Mitchell, J.K., Schlosser, F., and Dunnicliff, J., (1989), "Reinforced Soil Structures, Design and Construction Guide," 2 Vol., FHWA Report No. RD-89043.
Colman, B.D. and Knox, A.G., (1957), "The Interpretation of Creep Failure in Textile Fibers as a Rate Process," Textile Research Journal, PP. 393-399.
Coyle, H.M., and Reese, L.C., (1966), "Load Transfer for Axially Loaded Piles in Clay," ASCE, Journ. of Soil Mech. and Found. Div., Vol. 92, No.2, pp. 1-26.
Degoutte, G., and Mathieu, G., (1986), "Experimental Research of Friction Between Soil and Geomembranes or Geotextiles Using 30 x 30 cm2 Shearbox," 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 4, pp. 1251-1255.
El-Fermaoui, A., and Nowatzki, E., (1982), "Effect of Confining Pressure on Performance of Geotextiles in Soils," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 799-804.
Elias, V., (1979), "Friction in Reinforced Earth Utilizing Fine Grained Backfills," Internat. Conference of Soil Reinforcement, Paris, France, pp. 435-438.
Fowler, J., (1982), "Theoretical design Considerations for Fabric- Reinforced Embankments" 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 665-670.
Formazin, J., and Batereau, C., (1985), "The Shear Strength Behavior of Certain Materials on the Surface of Geotextiles," Proc. of the 11th. Internat. Conference on Soil Mechanics, San Francisco, Vol. 3, pp. 1773- 1775.
Forsyth, R. A., (1978), "Alternative Earth Reinforcement," Symposium on Earth Reinforcement, ASCE Convention, Pittsburgh, PA., pp. 358-362.
Gourc, J.P., Delmas, P., and Giroud, J.P., (1980), "Experiments on Soil Reinforcement with Geotextiles," ASCE Convention, Portland, Oregon.
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202
Gourc, J.P., Ratel, A., and Delmas, P., (1986), "Design of Fabric Retaining Walls: The Displacement Method," 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 3, pp. 289-294.
Greenwood, J.H., and Myles, B., (1986), "Creep and Stress Relaxation of Geotextiles," 3rd. Internat. Conference on Geotextiles, Vienna, Vol. 3, pp. 821-826.
Hausmann, M.R., and Vagneron, J.M., (1977), "Analysis of Soil-Fabric Interaction," Internat. Conference on the Use of Fabrics in Geotechnics, Paris, France, Vol. 3, pp. 139-144.
Hird, C.C., (1986), "Stability Charts for Reinforced Embankments on Soft Ground," Geotextiles and Geomembranes, 4, pp. 107-127.
Holtz, R.D., and Broms, B.B., (1977), "Walls Reinforced by Fabrics - Results of Model Tests," Internat. Conference on the Use of Fabrics in Geotechnics, Paris, France, Vol. 1, pp. 113-117.
Holtz, R.D., (1977), "Laboratory Studies of Reinforced Earth Using a Woven Polyester Fabric," Internat. Conference on the Use of Fabrics in Geotechnics, Paris, France, Vol.3, pp. 149-154.
Holtz, R.D., Tobin, W., and Burke, W .W., (1982), "Creep Characteristics and Stress-Strain Behavior of a Geotextile-Reinforced Sand," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp 805-809.
Ingold, T. S., and Templeman, J., (1979), "The Comparative Performance of Polymer Net Reinforcement," Internat. Conference on Soil Reinforcement, Paris, Vol .1
Ingold, T.S., (1982), "An Analytical Study of Geotextile Reinforced Embankments," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp 683-688.
Ingold, T.S., (1983), "Laboratory Pull-out Testing of Grid Reinforcement in Sand," Geotech. Testing Journal, 6(3), pp. 101-111.
Jaber, M., Mitchell, J.K., Shen, C.K., and Hua, Z.K., (1987), "Behavior of Reinforced Soil, Centrifuge Model Tests," FHWA Report No. DTFH61- 84-00073.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
203
Jacobsen, H. M., (1985), "Stability of Earth Structures Reinforced by Geotextiles," Proc. of the 11th. Internat. Conference on Soil Mechanics and Foundation Engineering, San Francisco, Vol. 3, pp. 1781-1785.
Jewell, R.A., (1980), "Some Effects of Reinforcement on the Mechanical Behavior of Soils," Ph.D. Thesis, Cambridge University, Cambridge, England.
Jewell, R.A., (1982), "A Limit Equilibrium Design Method for Reinforced Embankments on Soft Foundations," 2nd. Internat. Conference on Geotextiles, Las Vegas, Vol. 3, pp. 671-676.
Jewell, R.A., Milligan, G., Sarsby, R.W., and Dubois, D., (1984), "Interaction between Soil and Geogrids ", Proceedings, Polymer Grid Reinforcement, London, pp. 18-30.
Johnston, R.S., (1985), "Pull-out Testing of Tensar Geogrids," Thesis presented to University of California at Davis for M.Sc. Degree.
Juran, I., Knochenmus, G., Acar, Y.B., and Arman, A .,(1988), "Pull-out Response of Geotextiles and Geogrids (Synthesis of Available Experimental Data)," Proc. of Symp. on Geotextiles for Soil Improvement, ASCE, Geotech. Special Publication 18, pp. 92-111.
Juran, I., Guermazi, A., Chen, C.L., and Ider, M.H., (1988-b), "Modeling and Simulation of Load Transfer in Reinforced Soil: Part 1," Internat. Journal of Numer. and Analytical Methods in Geomechanics, Vol. 12, pp. 141-155.
Juran, I., Ider, H., Chen, C.L., and Guermazi, A., (1988-c), "Numerical Analysis of the Response of Reinforced Soils to Direct Shearing: Part 2," Internat. Journal of Numer. and Analytical Methods in Geomechanics, Vol. 12, pp. 157-171.
Juran, I., and Chen, C.L., (1988), "Soil-Geotextile Pull-out Interaction Properties, Testing and Interpretation," Transp. Research Board, 67th Annual Meeting, Washington D.C.
Juran, I., and Chen, C.L., (1989), "Strain Compatibility Design Method for Reinforced Earth Walls," Journal of Geotech. Engineering, ASCE, Vol. 115, No.4, pp. 435-456.
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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.
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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.
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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.
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Singh, A. and Mitchell, J.K., (1969), "Creep Potential and Creep Rupture of Soils," Proc. 7th. ICSMFE, Mexico, vol. 1.
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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.
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APPENDIX A
EQUIPMENT DETAILS & SPECIFICATIONS.
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209
- o
Figure A -1. Top View of the Pull-out Box
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210
H
H
Figure A-2. View of the Pull-out Box.
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Figure A-3. Detail of the Loading Frame.
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Figure A-4. Detail of the Loading Frame.
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K $ \
-4 -
Figure A-5. Detail of Load Cell Connections.
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nrjf do
Ola
214
4 - & d . a-.y.4 -
GD— IWCET
Figure A-6. Detail of Hopper.
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Si F £ £ =* 115 * s
V1 lii i .
5 !1
Figure A-7. Specifications of the LVDT's.
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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.
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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.
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
APPENDIX B
COMPACTION CONTROL AND DENSITY CALIBRATION.
218
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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).
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
APPENDIX C
PULL-OUT TEST RESULTS.
224
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
75 226
60
Test*2
45
■oMo►J
30
15
Unconfined Extension
Dl«pUc»ment (mm)
(a )
TEST #2
160
008
tooClMI
mm/secSlope mm/min)
IM 1600 •000 •M0 •000Time (oeo)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
TEST-A1060 -
/■Nai•Oao
0 SO 40 00 00 100 140 100120 100
Fact Dlsplacaaant M(a )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
/•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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
Load
(K
N/m
)
252
Unconfined Extension Tests
Geogrid: ConMed60Test: Con-U3
0 20 40 60 80 100D isplacem ent (mm)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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)
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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 )
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
APPENDIX D
FORMULATION OF SOIL-REINFORCEMENT
MODEL RELATIONSHIPS.
259
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
Reproduced with permission of the copyright owner. Further reproduction prohibited without permission.
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
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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.
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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:
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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.
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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.
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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.
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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
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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
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