OBSERVATIONS FROM LOAD TESTS ON GEOSYNTHETIC REINFORCED SOIL

OBSERVATIONS FROM LOAD TESTS ON

GEOSYNTHETIC REINFORCED SOIL

A THESIS SUBMITTED TO THE GRADUATE DIVISION OF THE

UNIVERSITY OF HAWAI‘I AT MĀNOA IN PARTIAL FULFILLMENT

OF THE REQUIREMENTS FOR THE DEGREE OF

MASTER OF SCIENCE

IN

CIVIL ENGINEERING

MAY 2014

By

Melia K. Iwamoto

Thesis Committee:

Phillip S. K. Ooi, Chair Person

Horst G. Brandes

Peter G. Nicholson

1

ACKNOWLEDGEMENTS

I would like to thank:

My advisor, Dr. Phillip Ooi, for the priceless guidance, instruction,

encouragement, and patience during my master’s research.

Mr. Mike Adams and Dr. Jennifer Nicks for allowing me to work on this project

and for the support during conference and journal submissions.

The TFHRC researchers and lab technicians for their quality work on the GRS

mini-piers.

Dr. Horst Brandes and Dr. Peter Nicholson for increasing my knowledge of

geotechnical engineering and for sitting in on my thesis committee.

Ms. Janis Kusatsu for the warm talks; Ms. Amy Fujishige for preparing my

paperwork for graduation.

Dr. Daniel Lipe for the countless hours of advice and emotional support.

The Lamb of God Church and Bible School for the prayers and encouragement.

And finally my family and friends, for keeping me grounded and making me

smile but most of all my mom, for giving me motivation when I’ve seemed to

have lost mine.

2

ABSTRACT

In lieu of conventional concrete cantilever retaining walls, geosynthetic reinforced

soil (GRS) is increasingly being used as bridge abutments to support single span bridges.

When backfilled with cohesionless material, GRS abutments have good drainage, large

bearing capacities, and great flexibility to withstand seismic loads and do not suffer from

the bump-at-the-end-of-the-bridge syndrome. GRS differs from geosynthetic

mechanically stabilized earth (GMSE) structures in many ways particularly with respect

to reinforcement spacing. The reinforcement spacing in GRS is closer (typically 8

inches) whereas the spacing in GMSE can be as large as 32 inches.

Four pairs of instrumented GRS square columns, also known as performance tests or

mini-piers, were load tested at the Turner-Fairbank Highway Research Center (TFHRC)

in McLean, Virginia from 2011 – 2012. Each pair consisted of identical reinforcement

strength to spacing ratio, backfill, and dimensions; however, one was loaded with a light-

weight dry-stacked friction-only concrete masonry unit (CMU) facing in place and the

other with the CMU removed prior to load testing. Moreover, each pair had different

reinforcement strength (Tf) and vertical spacing (Sv). The effects of the CMU facing and

varying Tf, Sv and Tf/Sv ratios on the GRS vertical capacity, load-settlement curves,

lateral deformation and lateral earth pressures during mini-pier construction and during

load testing were investigated. A method to deduce the composite shear strength

parameters using stress paths for pairs of these mini-piers was developed; the validity of

the postulate of zero volume change was investigated; and the appropriateness of using

fully softened versus peak backfill shear strength parameters to predict the bearing

capacity of a footing on a GRS abutment was studied.

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TABLE OF CONTENTS

Table of Contents .............................................................................................................. 3

List of Figures .................................................................................................................... 5

List of Tables ..................................................................................................................... 8

Introduction ................................................................................................................. 9

1.1 Introduction ............................................................................................................. 9

1.2 Motivation for Research ....................................................................................... 11

1.3 Thesis Overview ................................................................................................... 12

Literature Review ..................................................................................................... 14

2.1 Large-scale triaxial load tests and full-scale load tests on GRS ........................... 14

2.2 GRS Load Tests .................................................................................................... 15

2.2.1 Vegas Mini-Pier Experiments and the Postulate of Zero Volume Change

(Adams et al., 2002) ............................................................................................... 15

2.2.2 Large-Scale Unconfined Cylindrical Tests (Elton and Patawaran, 2005) .... 16

2.2.3 Mini Pier Experiments (Adams et al., 2007) ................................................ 20

2.2.4 GSGC Tests (Pham, 2009; Wu and Pham, 2013) ......................................... 22

2.2.5 TF tests and Validation Sessions (Nicks et al., 2013)................................... 24

2.2.6 University of Massachusetts at Amherst Load Tests (Mitchell 2002) .......... 27

2.3 Influence of Spacing versus Strength of Reinforcement on Performance of GRS 32

2.4 Bearing Capacity of a Footing on a GRS Abutment Wall .................................... 32

2.5 GRS versus GSME ............................................................................................... 36

Performance Test Program ...................................................................................... 38

3.1 Test Configuration ................................................................................................ 38

3.2 Backfill .................................................................................................................. 44

3.3 Geosynthetic Reinforcement ................................................................................. 47

3.4 Facing Elements .................................................................................................... 48

3.5 Loading System .................................................................................................... 48

3.6 Instrumentation ..................................................................................................... 50

3.6.1 Lateral and Vertical Deflection ..................................................................... 50

3.6.2 Lateral and Vertical Earth Pressure .............................................................. 51

3.6.3 Strain Gauges ................................................................................................ 52

Performance Test Results......................................................................................... 54

4.1 Ultimate Bearing Capacity .................................................................................... 54

4.2 Failure Plane ......................................................................................................... 59

4.3 Lateral Pressures ................................................................................................... 64

4

4.3.1 During Mini-Pier Construction ..................................................................... 64

4.3.2 During Load Testing ..................................................................................... 66

4.3.3 Lateral Earth Pressure Coefficients .............................................................. 68

4.4 Lateral Deformation .............................................................................................. 69

4.4.1 Postulate of Zero Volume Change ................................................................ 73

4.5 Fabric Strains ........................................................................................................ 78

4.5.1 Introduction ................................................................................................... 78

4.5.2 Strain Gauge Layout ..................................................................................... 79

4.5.3 Results ........................................................................................................... 80

4.6 Shear Strength Parameters of a GRS/GMSE ........................................................ 86

4.6.1 Construction of Stress Paths ......................................................................... 89

4.6.2 Strength Parameters ...................................................................................... 92

4.6.3 Discussion on the Shear Strength Parameters ............................................... 94

Use of Fully Softened Versus Peak Strengths to Predict the Ultimate Bearing

Capacity of Footings on GRS ......................................................................................... 96

5.1 Peak versus Fully Softened Strengths ................................................................... 96

5.1.1 Use of Peak Strengths to Predict Bearing Capacity ...................................... 99

5.1.2 Use of Fully Softened Strengths to Predict Bearing Capacity .................... 102

5.2 Summary ............................................................................................................. 103

Summary and Conclusions .................................................................................... 104

6.1 Summary ............................................................................................................. 104

6.2 Findings and Conclusions ................................................................................... 104

6.2.1 Bearing Capacity ......................................................................................... 104

6.2.2 Failure Plane: .............................................................................................. 105

6.2.3 Lateral Pressures: ........................................................................................ 105

6.2.4 Lateral Deformation: ................................................................................... 106

6.2.5 Postulate of Zero Volume Change .............................................................. 107

6.2.6 Fabric Strains .............................................................................................. 107

6.2.7 Shear Strength Parameters .......................................................................... 108

6.2.8 Fully Softened versus Peak Strengths ......................................................... 108

6.3 Key Findings ....................................................................................................... 109

6.4 Recommendations for Future Works .................................................................. 109

References ................................................................................................................ 111

5

LIST OF FIGURES

Figure 1-1 Typical cross-section of a GRS-IBS (Adams et al., 2011) ....................... 11

Figure 2-1 Schematic and photograph of Vegas Mini-Pier (Adams et al., 2002) ...... 16

Figure 2-2 Large-scale unconfined GRS load test (a) before loading and (b) after

failure (Elton and Patawaran, 2005) ......................................................... 19

Figure 2-3 Stress-strain curves of GRS with identical reinforcement strengths spaced

at 6 inches and 12 inches vertically (Elton and Patawaran, 2005)............ 19

Figure 2-4 Method to derive the (a) reinforced friction angle and (b) reinforced

cohesion of the GRS (Elton and Patawaran, 2005) ................................... 20

Figure 2-5 Schematic of Mini Pier dimensions and reinforcement spacing (in meters)

(Adams et al., 2007). ................................................................................. 21

Figure 2-6 (a) and (b) failures of widely spaced Mini Pier, MPs A and B; (c) and (d)

failures of closely spaced Mini Pier, MPs C and D (Adams et al., 2007) 22

Figure 2-7 (a) Application of grease to plexi glass within test frame to assure plane

strain conditions during loading (b) Latex membrane which enabled a

confining stress of 5 psi to be applied during loading (c) failed GSGC2

test specimen (Pham, 2009) ...................................................................... 24

Figure 2-8 Validation Session (VS) test set up (Nicks et al., 2013) ........................... 26

Figure 2-9 (a) Construction of TF-1 and (b) – (d) photographs test set up and loading

system of TF-1 through -3 (Nicks et al., 2013)......................................... 27

Figure 2-10 Variation of Ncq/2 with footing geometry and the stability factor. ........... 35

Figure 3-1 (a) Plan and profile schematic of TF-6; and (b) plan and profile schematic

of TF-7 ...................................................................................................... 40

Figure 3-2 (a) Schematic of TF-9; and (b) schematic of TF-10 ................................. 41

Figure 3-3 (a) Schematic of TF-12; and (b) schematic of TF-11 ............................... 42

Figure 3-4 (a) Schematic of TF-14; and (b) schematic of TF-13 ............................... 43

Figure 3-5 (a) Photograph of mini-pier test with CMU; (b) photograph of mini-pier

test without CMU; (c) photograph of TF-14............................................. 43

Figure 3-6 VDOT 21A grain size distribution ............................................................ 44

Figure 3-7 Compaction curve of VDOT 21A ............................................................. 45

Figure 3-8 Mohr-Coulomb failure envelope of GRS backfill based on large scale

direct shear tests ........................................................................................ 45

Figure 3-9 Schematic of CMU block.......................................................................... 48

Figure 3-10 Loading system ......................................................................................... 49

Figure 3-11 Schematics of deflection instrumentation for mini-piers (a) with CMU and

(b) without CMU....................................................................................... 51

Figure 3-12 Fatback cell mounted onto the CMU block. ............................................. 52

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Figure 3-13 Attachment of strain gauge on geotextile using the University of Colorado

at Denver attachment technique. ............................................................... 53

Figure 4-1 Ultimate capacity versus Tf/Sv .................................................................. 56

Figure 4-2 Ratio of ultimate capacities with and without CMUs versus reinforcement

spacing ...................................................................................................... 56


strength ...................................................................................................... 57

Figure 4-4 Load-settlement curves of the mini-pier load tests ................................... 58

Figure 4-5 Dimensionless form of the load-settlement curves of the mini-pier load

tests ........................................................................................................... 59

Figure 4-6 TF-11 at failure (qult,emp = 23.2 ksf) with Sv = 3-13/16 in and Tf = 1400

lb/ft (Nicks et al, 2013) ............................................................................. 60

Figure 4-7 TF-13 at failure (qult,emp = 13.0 ksf) with Sv = 11-1/4 in and Tf = 3600 lb/ft

(Nicks et al, 2013) ..................................................................................... 60

Figure 4-8 TF-10 at Failure (qult,emp = 10.33 ksf) with Sv = 15-1/4 in and Tf = 4800

lb/ft (Nicks et al, 2013) ............................................................................. 61

Figure 4-9 (a) – (i) Rupture pattern for geotextiles in TF-6; (j) Schematic of tears in

geotextiles in TF-6. ................................................................................... 63

Figure 4-10 Measured lateral pressures at Fatback cell during construction of Mini-

Piers........................................................................................................... 66

Figure 4-11 Measured lateral pressures at Fatback cell location during load testing of

TF-6, -9, -12, and -14. ............................................................................... 67

Figure 4-12 Lateral earth pressure coefficients versus dimensionless lateral movement.

................................................................................................................... 69

Figure 4-13 Lateral deformation profiles of GRS at loads ≈ 83% of ultimate load. .... 71

Figure 4-14 Lateral displacement of TF-6 with increasing applied load (with CMU)

Tf/Sv = 7.55 ksf. ........................................................................................ 72

Figure 4-15 Lateral deformation of TF-7 with increasing applied load (without CMU)

Tf/Sv = 7.55 ksf. ........................................................................................ 72

Figure 4-16 Schematic of assumed deformed mass for TF-11 at applied load = 23 ksf

(drawn to scale) as assumed by Adams, et al. (2002) and TF tests in (a)

profile view and (b) plan view. ................................................................. 74

Figure 4-17 Volumetric strain versus vertical applied load .......................................... 76

Figure 4-18 Volumetric strain due to service loads (4 ksf). ......................................... 77

Figure 4-19 Volume gained versus volume lost ........................................................... 78

Figure 4-20 Strain measured using strain gauges and POTs versus distance from the

center of the GRS mass for (a) Geotextile 7 (b) Geotextile 5 and (c)

Geotextile 3 for TF-6. ............................................................................... 83

7



Geotextile 6 for TF-11. ............................................................................. 84



Geotextile 8 for TF-12. ............................................................................. 85

Figure 4-23 Stress paths during load testing of TF-13 (without CMU) and TF-14 (with

CMU) with Tf/Sv = 3.84 ksf. ..................................................................... 87

Figure 4-24 Stress paths during load testing of TF-9 (with CMU) and TF-10 (without

CMU) with Tf/Sv = 3.78 ksf. ..................................................................... 87


CMU) with Tf/Sv = 4.41 ksf. ..................................................................... 88


CMU) with Tf/Sv = 7.55 ksf. ..................................................................... 88

Figure 4-27 Free-body diagram of vertically loaded GRS ........................................... 91

Figure 4-28 (a) and (b) Photographic evidence of when CMU blocks lurched forward

at failure of TF-6; (c) and (d) Photographic evidence of when CMU blocks

lurched forward at failure of TF-12. ......................................................... 93

Figure 4-29 Dimensionless forms of cohesion and friction angle versus dimensionless

for of Tf/Sv ratio ........................................................................................ 95

Figure 5-1 (a) Large scale direct shear device (LSDS) (b) Large scale triaxial device

(LSTX) ...................................................................................................... 98

Figure 5-2 Typical stress-strain curve of an overconsolidated soil ............................ 98

Figure 5-3 Histogram and normal distribution of the bias using peak strengths ...... 101

Figure 5-4 Measured versus predicted capacities using peak strengths ................... 101

Figure 5-5 Histogram and normal distribution of the bias using fully softened

strengths .................................................................................................. 102

Figure 5-6 Measured versus predicted capacities using fully softened strength ...... 103

8

LIST OF TABLES

Table 2-1 Summary of test variables and results from large-scale unconfined

cylindrical GRS load tests (Elton and Patawaran, 2005) .......................... 18

Table 2-2 Test variables of Mini Pier experiments (Adams et al., 2007) ................. 21

Table 2-3 Test variables of GSGC tests (Pham, 2009) ............................................. 24

Table 2-4 Summary of test variables for TF and VS tests (Nicks et al., 2013)......... 25

Table 2-5 Influence of spacing versus tensile strength on the bearing capacity of a

footing on a GRS ...................................................................................... 32

Table 3-1 Mini-pier load tests conducted at TFHRC ................................................ 39

Table 3-2 Properties of geotextiles used in the mini-pier load or performance tests

conducted at TFHRC. ............................................................................... 48

Table 3-3 Instruments deployed in TF tests. ............................................................. 50

Table 4-1 Ultimate bearing capacity and strain at failure of mini-pier load tests ..... 55

Table 4-2 Performance tests dates and Fatback cell temperatures ............................ 65

Table 4-3 Summary of strain gages in performance tests ......................................... 79

Table 4-4 Shear strength parameters of GRS ............................................................ 93

Table 5-1 Test parameters of GRS performance tests selected from literature ......... 97

Table 5-2 Predicted and measured ultimate bearing capacity of GRS load tests using

fully softened versus peak strengths ....................................................... 100

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INTRODUCTION

1.1 Introduction

Geosynthetic reinforced soil (GRS) is defined as closely-spaced (≤ 12 inches;

typically ≈ 8 inches) alternating layers of geosynthetic reinforcement and compacted

granular fill material (Adams, et al., 2011). GRS has been used for a variety of

geotechnical applications but has recently been promoted by the Federal Highway

Administration (FHWA) for use as abutments for single span steel or concrete bridges in

their Everyday Counts Initiative, which is focused on accelerating implementation of

proven, market-ready technologies. GRS-IBS, where IBS stands for Integrated Bridge

System, consists of a reinforced soil foundation (RSF), a GRS abutment and a GRS

integrated approach (Figure 1-1). The RSF consists of granular fill compacted and

encapsulated in geotextile. The RSF provides embedment and increases the bearing width

and capacity of the GRS abutment. The GRS abutment provides load-bearing support for

the bridge, which is placed directly on the abutment. GRS is also used to construct the

integrated approach adjacent to the superstructure. GRS-IBS has the following

advantages:

It is a fast and cost-effective method of bridge support which eliminates the need

for cast-in-place reinforced concrete abutments traditionally supported on deep

foundations.

Quality compaction control can be realized since closely-spaced geosynthetics

ensure backfill is placed in thin lifts.

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Catastrophic collapse was not observed in numerous load tests carried out to

failure; GRS abutments behave in a ductile fashion.

It can be built in variable weather with common labor, materials and equipment,

and can be easily modified in the field.

The “bump-at-the-end-of-the-bridge” problem caused by differential settlement

between the bridge abutment and the approach roadway is alleviated. This is

made possible by eliminating deep foundations, by using GRS to construct the

integrated approach, and by limiting its use to short, single-span integral bridge

systems.

It enjoys all the advantages associated with an integral abutment bridge.

A very flexible system, it is amenable to differential settlement and seismic

loading.

11

Figure 1-1 Typical cross-section of a GRS-IBS (Adams et al., 2011)

GRS differs from geosynthetic mechanically stabilized earth (GMSE) structures in

many ways particularly with respect to the maximum reinforcement spacing. The

reinforcement spacing in GRS is closer whereas the spacing in GMSE can be as large as

32 inches. GMSE and GRS are also designed differently. Therefore, it is of interest to

the geotechnical engineer to understand differences in their behavior upon loading.

1.2 Motivation for Research

Because GRS-IBS is load bearing, its capacity is an important design consideration.

Many large-scale load tests to determine the bearing capacity of GRS has been performed

and are available in the literature. Recently, an extensive series of load tests on

GRS/GMSE mini-piers or square columns were performed at FHWA's Turner-Fairbank

12

Highway Research Center (TFHRC) in McLean, Virginia with and without concrete

masonry unit (CMU) facing frictionally connected to the reinforcement (Nicks et al.

2013). The tests were instrumented to provide valuable information on the internal

behavior of the GRS and can be used to understand the influence of the facing,

reinforcement strength (Tf) and reinforcement spacing (Sv) on the bearing capacity of a

footing on the GRS.

In light of this, the objectives of this research were to: (1) investigate the effects of

modular block facing with frictional connection and varying Tf/Sv (reinforcement

strength to vertical spacing) ratios on the GRS vertical capacity, load-settlement curves,

lateral deformation and lateral earth pressures during mini-pier construction and during

load testing, (2) derive the composite shear strength parameters of the GRS/GMSE; (3)

examine the validity of the postulate of zero volume change (Adams et al., 2002) during

loading of GRS, which forms the basis for estimating lateral deformations of GRS walls

and (4) assess the appropriateness of using the soil’s fully softened versus peak shear

strengths to predict the bearing capacity of footings on GRS.

1.3 Thesis Overview

An outline of the thesis is provided below:

Chapter 2 Findings from a literature review on GRS load tests, method of

estimating the bearing capacity of a footing on a GRS wall, and

differences between GRS and GMSE

Chapter 3 Description and test set up of the TFHRC performance tests

Chapter 4 Presentation of results from the TFHRC performance tests

13

Chapter 5 Study on the appropriateness of using the soil’s fully softened versus

peak shear strengths to predict the bearing capacity of footings on GRS

Chapter 6 Summary and conclusions

14

LITERATURE REVIEW

2.1 Large-scale triaxial load tests and full-scale load tests on GRS

Many large-scale triaxial tests have been performed on GRS reinforced soil (Chen

et al., 2000; Lee, 2000; Ketchart and Wu, 2001; Holtz and Lee, 2002; Zhang et al., 2006;

Ziegler, 2008; etc.). These tests are useful for shedding light on the influence of

reinforcement spacing and ultimate tensile strength on the strength and performance of

GRS. However, the U.S. Army Corps of Engineers (Jewell 1993) suggested that the test

specimen must be at least 6 times larger than the maximum particle size and at least 15

times larger than the average particle size of the backfill to eliminate particle size effects

on the test specimen.

Another concern with laboratory triaxial tests on GRS is the limit on reinforcement

spacing. The spacing in the field cannot be mimicked with triaxial size specimens.

Many full-scale GRS abutment experiments have been performed (Adams et al.,

1997; Gotteland et al., 1997; Ketchart and Wu, 1997; Tatsuoka et al., 1997; Wu et al.,

2006; Yoo and Kim, 2008). These tests demonstrated the reliability of GRS walls

subjected to typical working loads, but they were not loaded to failure. Also, a series of

large-scale plane strain load tests were performed on geogrid reinforced soil structures at

the Royal Military College of Canada (Bathurst et al., 1988; Bathurst and Benjamin,

1990; Bathurst et al., 2000; Burges, 1999; Vlachopoulos, 2000; Saunders, 2001; Leblanc,

2002; Reeves, 2003; Nelson, 2005). However, these tests will not be discussed further

because: (1) the tests involve geogrid-reinforced soils with large rectangular panels as

facing. These panels are significantly larger than the modular block facing typically

15

associated with GRS. In some tests, the panels were stacked horizontally while in others,

they stood vertically side-by-side; (2) the connections between the facing panels and the

reinforcement were not frictional. In fact, the connection strengths are believed to be

much higher than the frictional connection in a GRS; (3) the reinforcement spacings were

higher than 12 inches in most tests; (4) some of the load tests on the geogrid walls were

not carried out to failure; and (5) the soil’s fully softened strengths could not be discerned

from the publications.

Large-scale GRS load tests carried out to failure are expensive to perform and are

primarily research-based. Hence, there are only a few available, a summary of which is

presented in the sections below.

2.2 GRS Load Tests

2.2.1 Vegas Mini-Pier Experiments and the Postulate of Zero Volume Change (Adams

et al., 2002)

In 2000, the FHWA began its first series of large-scale GRS load tests with the

Vegas Mini-Pier (Adams et al., 2002). This test was not loaded to failure due to time

constraints and stroke limitations. However, it is discussed because this publication

introduced the postulate of zero volume change.

This test was performed on a square column (3.5 ft x 3.5 ft x 8 ft high) with

segmental retaining wall (SRW) blocks as facing (Figure 2-1). The geotextile was a

woven polypropylene having an ultimate tensile strength (Tf) of 2400 lb/ft spaced (Sv)

every 6 inches giving a Tf/Sv ratio of 4800 psf. Bearing bed reinforcements were

16

provided in the top two rows (Sv = 3 in). The soil was a poorly graded-silty gravel

backfill (GP-GM).

This test was instrumented to measure applied vertical load, settlement, and lateral

deformation. When loaded, Adams et al. (2002) found that the lateral expansion was

equal to the vertical compression of the GRS. This along with other similar load tests

gave rise to the postulate of zero volume change, which can be used to predict the

maximum lateral displacement of a GRS wall based on its settlement, and vice versa. An

objective of this study is to validate the postulate of zero volume change (discussed in

greater detail in Section 4.4.1).

Figure 2-1 Schematic and photograph of Vegas Mini-Pier (Adams et al., 2002)

2.2.2 Large-Scale Unconfined Cylindrical Tests (Elton and Patawaran, 2005)

Elton and Patawaran (2005) performed unconfined load tests on 7 large-scale

cylindrical columns (2.5ft diameter x 5 ft high) reinforced with nonwoven polypropylene

geotextiles of varying tensile strengths at a constant vertical spacing of 6 inches (Table 2-

1 and Figure 2-2). The backfill was a poorly graded sand (SP) with less than 5% fines.

17

An initial test was performed with a reinforcing spacing of 12 inches. Upon conducting a

post-mortem of this test column, it was found that the reinforcement strength was not

mobilized (no tears in the geotextile). The stress-strain curves of two GRS columns

utilizing the same geotextile but one having a 12 inch spacing and the other 6 inch

spacing are shown in Figure 2-3. From these two tests, Elton and Patawaran (2005)

noted that the reinforcement strength of the 12-inch-spaced GRS column was not

mobilized. Therefore, subsequent tests were spaced 6 inches apart. These tests were

instrumented to measure vertical applied load, lateral deformation and settlement.

These tests were run on asymmetrically loaded cylindrical columns, which is very

different from the plain strain condition of a footing on a GRS abutment. Nevertheless,

they provide valuable information to this study.

The peak cohesion and friction angle of the backfill determined from direct shear

tests were 0.6 ksf and 40°, respectively. Elton and Patawaran (2005) estimated the

friction angle of the composite, ϕreinf, as the angle of a line that is drawn through the soil’s

cohesion = 0.6 ksf and that intersects the reinforced soil’s Mohr circle at failure (Figure

2-4a). The cohesion of the composite was calculated by drawing a 40° line tangent to the

reinforced soil’s Mohr circle at failure. Where this line crosses the y-axis was taken as

creinf (Figure 2-4b).

The method to determine the shear strength parameters of the GRS composite

follow the concept of apparent cohesion (Yang, 1972) to derive creinf and the concept of

apparent confining pressure (Schlosser and Long, 1972) to derive ϕreinf. However, these

methods were incorrectly executed. First, the cohesion that should be used to determine

18

the reinforced friction angle is not the backfill cohesion but rather a function of the

backfill friction angle, reinforcement spacing and strength which has been shown by

Yang (1972) to be:

𝑐𝑅′ =

∆𝜎3

2√𝐾𝑝 + 𝑐 =

𝑇𝑓√𝐾𝑝

2𝑆𝑣+ 𝑐 (2-1)

where Kp is the Rankine passive earth pressure coefficient = tan2(45°+ϕ/2), Tf and Sv are

the reinforcement strength and spacing, and c is the cohesion of the backfill. One note on

Yang’s (1972) concept of apparent cohesion is that the assumption that the increased

apparent confining pressure is equal to Tf/Sv (implying that increasing Tf has the same

effect as a proportional decrease in Sv) has been proven to be untrue (see Section 2.3 for

evidence on why this concept is incorrect).

Table 2-1 Summary of test variables and results from large-scale unconfined cylindrical

GRS load tests (Elton and Patawaran, 2005)

Test

USCS

Symbol

Tf

(lb/ft)

Sv

(in)

Tf/Sv

(ksf)

Facing

Type

Boundary

Condition H/B1

qult,emp2

(ksf)

Elton1 SP 612 6.0 1.22 None Cylindrical Column 2 4.81







Notes: 1. H = specimen height and B = specimen diameter

2. qult,emp = measured ultimate bearing capacity of the GRS test

19

(a) (b)

Figure 2-2 Large-scale unconfined GRS load test (a) before loading and (b) after failure

(Elton and Patawaran, 2005)

Figure 2-3 Stress-strain curves of GRS with identical reinforcement strengths spaced at 6

inches and 12 inches vertically (Elton and Patawaran, 2005)

20

(a) (b)

Figure 2-4 Method to derive the (a) reinforced friction angle and (b) reinforced cohesion

of the GRS (Elton and Patawaran, 2005)

2.2.3 Mini Pier Experiments (Adams et al., 2007)

Adams et al. (2007) performed five large-scale unconfined GRS experiments,

referred to as Mini Piers (MP) in 1997 at the TFHRC. The MPs (39 in. x 39 in. x 76.25

in. high) utilized woven geotextiles of varying spacing and tensile strengths and a well

graded silty gravel (GW-GM) backfill as specified by the USCS classification system

(Table 2-2, Figure 2-5). They were constructed using CMU blocks that were removed

prior to loading in order to negate the effect of facing on the performance of the GRS and

to observe soil deformation. The first mini-pier, MPA, was constructed with 3 geotextile

layers with the top and bottom reinforcement layers placed two CMU courses below the

top and above the bottom, respectively, and the middle reinforcement spaced three CMU

courses apart (Figure 2-5; dimensions given in meters Adams et al., 2007). Results from

these tests suggested that the frequency of spacing (MPC has higher capacity than MPA

21

or MPB) has a greater influence on the bearing capacity of the GRS than reinforcement

strength (Figure 2-6).

Table 2-2 Test variables of Mini Pier experiments (Adams et al., 2007)

Test

USCS

Symbol

Tf

(lb/ft)

Sv

(in)

Tf/Sv

(ksf)

Facing

Type

Boundary

Conditions H/B1

qult,emp

(ksf)

MP NR GW-GM None None None None Square Column 2 0.90

MPA GW-GM 4800

15.25 –

22.875

3.78 –

2.52 None Square Column 2 4.70

MPB GW-GM 4800 15.25 3.78 None Square Column 2 3.55

MPC GW-GM 1400 7.625 2.20 None Square Column 2 9.61

MPD GW-GM 4800 7.625 7.55 None Square Column 2 -2

Notes: 1. H/B = Mini Pier height/width ratio

2. Mini Pier did not fail due to insufficient stroke

Figure 2-5 Schematic of Mini Pier dimensions and reinforcement spacing (in meters)

(Adams et al., 2007)

22

(a) (b)

(c) (d)

Figure 2-6 (a) and (b) failures of widely spaced Mini Pier, MPs A and B; (c) and (d)

failures of closely spaced Mini Pier, MPs C and D (Adams et al., 2007)

2.2.4 GSGC Tests (Pham, 2009; Wu and Pham, 2013)

Pham (2009) performed five large-scale GRS load tests, referred to as “Generic

Soil-Geosynthetic Composite” or GSGC tests. The GSGC tests were set up in plane

strain to more closely resemble a footing on a GRS abutment by using greased plexi-glass

23

at two opposite ends. The plexi-glass ends are restricted from lateral strain in the out-of-

plane direction. Photographs of the application of grease to the plexi-glass within the test

frame, the latex membrane which enabled a confining pressure of 5 psi to be applied to

the GRS, and a failed GSGC specimen are shown in Figures 2-7(a) – (c). Table 2-3

summarizes details of tests GSGC1 through GSGC5. GSGC1 was unreinforced while

GSGC 2 through 5 were reinforced. GSGC1 through GSGC4 were subjected to a

confining stress of 5 psi while GSGC5 had no applied confining pressure.

Composite strength parameters were obtained by applying Schlosser and Long’s

method (1972) and assuming ϕreinf = ϕbackfill. The apparent cohesion was calculated using

Equation 2-1. Pham (2009) observed that a GRS with a reinforcement strength Tf at

spacing Sv will not have the same capacity as a GRS with a reinforcement strength 2Tf at

spacing 2Sv. Realizing that the spacing has a greater influence than the reinforcement

strength, he developed Equation 2-2 to estimate the composite cohesion of a GRS.

Equation 2-2 includes a W-factor that amplifies the contribution of spacing to GRS

capacity.

𝑐𝑅 =∆𝜎3

2√𝐾𝑝 + 𝑐 = 𝑊

𝑇𝑓

2𝑆𝑣√𝐾𝑝 + 𝑐 = [0.7

(𝑆𝑣

6𝑑𝑚𝑎𝑥)]

𝑇𝑓

2𝑆𝑣√𝐾𝑝 + 𝑐 (2-2)

where Δ3 = increased confining pressure in a GRS mass due to the reinforcement, Kp =

Rankine passive earth pressure coefficient, c = backfill cohesion, dmax = maximum

particle size of the soil, Sv = reinforcement spacing, and Tf = reinforcement tensile

strength. The friction angle of the GRS was assumed to be equal to that for the

unreinforced soil. Pham’s W-factor was used to develop a model to predict the bearing

24

capacity of a footing on a GRS abutment, the derivation of which is discussed in Section

2.3. This model is included in the FHWA’s GRS-IBS Interim Implementation Guide and

Synthesis Report [FHWA, 2011(a) and 2011(b)] as a method to predict the bearing

capacity of a footing on a GRS abutment.

Table 2-3 Test variables of GSGC tests (Pham, 2009)

Test

No.

USCS

Symbol

Tf

(lb/ft)

Sv

(in)

Facing

type

Boundary

Conditions

H1

(in)

B2

(in)

L3

(in)

qult,emp

(ksf)

GSGC1 GW-GM - - 4 Plain Strain 76.25 57 47 16.09

GSGC2 GW-GM 4800 7.63 4 Plain Strain 76.35 54 46.75 70.95

GSGC3 GW-GM 9600 15.25 4 Plain Strain 76.35 53 46.75 42.58

GSGC4 GW-GM 4800 15.25 4 Plain Strain 76.30 58.75 46.75 37.25

GSGC5 GW-GM 4800 7.63 None Plain Strain 76.35 49 46.75 42.45

Note: 1. H = GSGC height (in)

2. B = GSGC width (in)

3. L = GSGC length (in)

4. No facing was used. Instead, a confining pressure = 5 psi was applied with a

latex rubber membrane wrapped around the GRS

(a) (b) (c)

Figure 2-7 (a) Application of grease to plexi glass within test frame to assure plane strain

conditions during loading (b) Latex membrane which enabled a confining stress of 5 psi

to be applied during loading (c) failed GSGC2 test specimen (Pham, 2009)

2.2.5 TF tests and Validation Sessions (Nicks et al., 2013)

The FHWA reported 19 load tests on square GRS columns (Nicks et al., 2013).

Five load tests were performed at the Defiance County, Ohio highway maintenance

25

facility and are referred to as Validation Sessions or “VS” (Figure 2-8). Fourteen load

tests were performed at the TFHRC in McLean, Virginia and are denoted as “TF”. Table

2-4 contains a summary of the test variables. The VS tests and TF-8 included two

courses of bearing bed reinforcement placed at the top of the mini pier. VS-1, -2, -3, -5

and TF-1 utilized open-graded backfill (GP), VS-4 poorly graded sand (SP), and the

remaining tests well-graded backfill (GW-GM). Difficulties in testing rendered the

results from VS-3, -4, TF-4 and -5 unreliable. TF-8 is not considered herein because it is

similar to TF-7 except that it has the bearing bed reinforcement.

Table 2-4 Summary of test variables for TF and VS tests (Nicks et al., 2013)

Test

No.

USCS

Symbol

Tf

(lb/ft)

Sv

(in)

Tf/Sv

(lb/ft2)

Facing

type

Boundary

Conditions H/B1

qult,emp

(ksf)

VS-1 GP 4800 7.63 7600 CMU Square Column 2 23.32


VS-3 GP 4800 7.63 7600 CMU Square Column 2 -

VS-4 SP 4800 7.63 7600 CMU Square Column 2 -


TF-1 GP 2400 7.63 3800 CMU Square Column 2 20.50

TF-2 GW-GM 2400 7.63 3800 CMU Square Column 2 25.26

TF-3 GW-GM 2400 7.63 3800 None Square Column 2 17.49

TF-4 GW-GM 4800 7.63 7600 None Square Column 2 -











Note: 1. H/B = GRS mini-pier height/width ratio

26

Figure 2-8 Validation Session (VS) test set up (Nicks et al., 2013)

Figures 2-9(a) shows the backfill and construction of TF-1. Figures 2-9(b)

through (d) are photographs of TF-1, -2, and -3, respectively, just before loading. Notice

that the loading system used for these tests are different than as described in Chapter 3 for

TF-6 through -14. For these tests, four Enerpac hydraulic jacks applied loads at the ends

of two steel beams which transferred the load to the concrete footing. Uneven loading to

the GRS inspired the transition to the new loading system employed in the remaining

tests. Note that TF-9 and -10 did not meet minimum GRS spacing requirements (≤ 12

inches).

27

(b)

(c) (d)

Figure 2-9 (a) Construction of TF-1 and (b) – (d) photographs test set up and loading

system of TF-1 through -3 (Nicks et al., 2013)

2.2.6 University of Massachusetts at Amherst Load Tests (Mitchell 2002)

Mitchell (2002) measured lateral pressures on the concrete masonry unit (CMU)

facing of four GRS mini-piers (Figure 2-9). The reinforcement (Amoco 2066 - wide

width tensile strength = 2100 lb/ft) spacings were 24 in, 16 in, 8 in and 32 in in piers 1, 2,

28

3 and 4, respectively. The facing consisted of 99-lb split face keystone retaining wall

blocks that were 18 in long, 5 in wide and 8 in deep with a flange protruding to the inside

of the piers. Corner blocks with slightly different weight and dimensions were used.

Known as trap rock gravel (32.1 % gravel, 56.5% sand and 11.5% fines, Cu = 40,

Cc = 3.2, Gs = 2.93, Standard Proctor d max = 144 pcf, wopt = 8.75%, Modified Proctor d

max = 150 pcf, wopt = 6.5%), the backfill had shear strength parameters of = 36˚ and c =

637 psf based on a failure relative displacement of 10% of the shear box length as

measured in a 12-inch x 12-inch direct shear box. The shear stress-displacement curves

appear to still be strain hardening at the interpreted failure relative displacement. Hence

these tests were not considered in this research.

(a)

29

(b)

Figure 2-9 GRS mini-pier (a) elevation (b) plan (after Mitchell, 2002)

Geokon vibrating wire pressure transducers were placed at 5 elevations (1 ft, 4.3 ft, 7 ft,

9.7 ft and 13 ft above base) along the GRS wall height. The measured lateral pressures

are shown in Figure 2-10. These pressure readings were zeroed at zero load and

represent only the increase in the lateral pressures during the load test. Mitchell (2002)

noted that with a decrease in reinforcement spacing, the lateral pressures increased and

became more equally distributed over the wall height.

The actual measured lateral pressures at zero load were provided in Appendix C

of Mitchell's (2002) report. Unfortunately, they are mostly negative with a few

exceptions. When and how the zero readings were obtained could not be discerned from

Mitchell's (2002) report.

30

Figure 2-10 Measured increase in lateral pressures on the facing of GRS mini-piers

during the load test (after Mitchell, 2002)

31

32

2.3 Influence of Spacing versus Strength of Reinforcement on Performance of GRS

Large-scale GRS load tests indicate that reinforcement spacing, Sv, has a greater

influence on the bearing capacity than the reinforcement ultimate tensile strength, Tf

(Adams et al., 2002; Elton and Patawaran, 2005; Adams et al., 2007; Pham, 2009; and

Nicks et al., 2013). Table 2-5 presents the ratio of measured ultimate bearing capacities

when the spacing is doubled versus when the reinforcement strength is doubled. From

this table, it can be seen that the increase in bearing capacity due to doubling the

reinforcement spacing is 2 to 67% higher than when doubling the reinforcement strength.

Table 2-5 Influence of spacing versus tensile strength on the bearing capacity of a footing

on a GRS

Test

No.

Tf

(lb/ft)

Sv

(in)

qult,emp

(ksf)

qult,emp, Sv/

qult,emp,2Sv

qult,emp,2Tf/

qult,emp,Tf

Percent

Difference

GSGC2 4800 7.63 70.96 1.90

67% GSGC4 4800 15.25 37.25 1.14

GSGC3 9600 15.25 42.57

Elton2 612 12 2.70 1.78

2% Elton1 612 6 4.81 1.75

Elton5 1272 6 8.40

TF-9 4800 15.25 22.31 1.96

13% TF-6 4800 7.63 43.76 1.73

TF-2 2400 7.63 25.26

2.4 Bearing Capacity of a Footing on a GRS Abutment Wall

Because GRS-IBS is load bearing, its capacity is an important design consideration.

Many large-scale load tests to determine the bearing capacity of GRS have been

performed (Adams, 1997; Adams et al., 2007; Elton and Patawaran, 2005; Mitchell,

2002; Pham, 2009).

33

Pham (2009) derived the bearing capacity of a footing on a GRS abutment wall

(qult) as follows:

𝑞ult = (𝜎h + 𝑊𝑇𝑓

𝑆𝑣) 𝐾𝑝 + 2𝑐√𝐾𝑝 (2-3)

where σh is the lateral stress, Tf and Sv are the reinforcement strength and spacing,

respectively, c is the soil cohesion, Kp is the Rankine passive earth pressure coefficient,

defined as

𝐾𝑝 =1+sin 𝜙

1−sin 𝜙 (2-4)

is the soil friction angle. W is a dimensionless factor that amplifies the contribution of

Sv to the GRS capacity, and was semi-empirically derived as

𝑊 = 0.7𝑆𝑣

6𝑑𝑚𝑎𝑥 (2-5)

where dmax is the maximum particle size of the GRS backfill. Note that the 0.7 factor was

theoretically derived using the concept of “average stresses” proposed by Ketchart and

Wu (2001) while the exponent was empirically derived. For details on this derivation,

refer to Pham (2009).

For a GRS wall with dry stacked modular block facing, σh = lateral stress exerted

by the facing on the GRS mass, defined by Pham (2009) as

𝜎ℎ = 𝛾𝑏𝑙𝐷 tan 𝛿 (2-6)

where γbl = bulk unit weight of facing block = weight of block/volume of block assuming

it is not hollow, D = depth of facing block perpendicular to the wall face and δ = friction

angle between geosynthetic reinforcement and the top and bottom surface of the facing

block.

34

Applicability of the Bearing Capacity Equation

Considering that most of the load tests in the database were performed on GRS

columns (mostly square with some circular in plan) while a bridge footing resting on an

abutment more resembles a plane strain (PS) condition, the relationship between the

column tests and that of a strip footing loading the top of a GRS wall is of interest.

Assume that the strength of a GRS column can be represented by the Mohr-Coulomb

equation as follows:

𝜏 = 𝑐𝐺𝑅𝑆 + 𝜎𝑡𝑎𝑛𝜙𝐺𝑅𝑆 (2-7)

where = shear strength, = applied normal stress, cGRS and GRS = cohesion and friction

angle of the GRS composite, respectively. In an unconfined compression load or

Performance Test (PT), where the facing has been removed, the ultimate capacity of the

GRS column (qult,PT) can be expressed as

𝑞𝑢𝑙𝑡,𝑃𝑇 = 2𝑐𝐺𝑅𝑆 (2-8)

For the PS condition, the bearing capacity of a footing supporting the bridge

superstructure can be estimated using Meyerhof’s (1957) solution for a rough strip

bearing on top of a slope

𝑞𝑢𝑙𝑡,𝑃𝑆 = 𝑐𝐺𝑅𝑆𝑁𝑐𝑞 + 0.5𝛾𝐺𝑅𝑆𝑏𝑁𝛾𝑞 (2-9)

where qult,PS = ultimate capacity of strip footing under PS conditions, GRS = unit weight

of the GRS backfill, b = footing width, and Ncq and Nq = Meyerhof’s (1957) bearing

capacity factors for a strip footing with a rough base. Nq approaches zero when the slope

angle is 90 for a GRS abutment wall; thus Equation (2-9) reduces to

35

𝑞𝑢𝑙𝑡,𝑃𝑆 = 𝑐𝐺𝑅𝑆𝑁𝑐𝑞 (2-10)

Dividing Equation (2-10) by (6), the ratio of the bearing capacity of a strip footing

on top of a GRS abutment to that of an unconfined GRS column can be estimated as

𝑞𝑢𝑙𝑡,𝑃𝑆

𝑞𝑢𝑙𝑡,𝑃𝑇=

𝑁𝑐𝑞

2 (2-11)

For a surface footing on top of a vertical GRS abutment, the value of Ncq varies

with the footing offset from the edge of the wall face, a, wall height, H, footing width, b,

and stability factor, 𝑁𝑠 =𝛾𝐺𝑅𝑆𝐻

𝑐𝐺𝑅𝑆, as shown in Figure 2-10.

Figure 2-10 Variation of Ncq/2 with footing geometry and the stability factor

cGRS can be obtained from laboratory or numerical experiments. For example, Pham

(2009) conducted a series of plane strain load tests on 6.36-ft-high GRS that can be used

to derive a cohesion value for the GRS. Two of the tests (GSGC2 and 5 in Table 2-3)

were identical in every respect (Tensile strength of reinforcement = 4800 lb/ft,

0

1

2

3

0 1 2 3 4 5

qu

lt,P

S/q

ult

,PT

a/H

Ns = 0

Ns = 2

Ns = 4

36

Reinforcement spacing = 7.625 in, Backfill c = 1562 psf and = 50º) except for the

confining stresses (0 in GSGC5 and 5 psi in GSGC2). The corresponding failure stresses

were 42.45 and 70.95 ksf for the 0 and 5 psi confining stresses, respectively. The

resulting shear strength parameters for the GRS are cGRS = 3343 psf and GRS = 72. The

corresponding stability factor H/c ≈ 0.29. Based on this stability factor, the ratio of

plane strain capacity for a typical GRS abutment with a typical set-back a = 7.63 in and H

varying from 10 ft to 33 ft (i.e. a/H = 0.02 to 0.07) to column (PT) capacity is close to

unity. Therefore, the column PT is fairly representative of an in-service PS condition for

well-graded gravels in this case.

2.5 GRS versus GSME

GRS differs from geosynthetic mechanically stabilized earth (GMSE) structures in

many ways particularly with respect to the maximum reinforcement spacing. The

reinforcement spacing in GRS is closer whereas the spacing in GMSE can be as large as

32 inches. GMSE is typically designed using the Simplified Method (SM), which was

developed from the Tieback Wedge Method (Bell et al., 1975). The SM applies to both

extensible and inextensible reinforcements. For the latter, the SM assumes a bi-linear

failure plane in accordance with the Coherent Gravity Method (Anderson et al., 2010)

while the linear Rankine active failure wedge is assumed for the former. Regardless of

reinforcement type, the SM models the GMSE reinforcements as tieback elements that

provide tensile resistance to the driving forces. Because the reinforcement layers in a

GRS are much closer, they increase the confinement, reduce lateral soil displacements

and limit soil dilation in addition to providing tensile resistance. Consequently, the

37

vertical and lateral displacements in a GRS are smaller, the bearing capacity of footings

on a GRS is larger and the lateral thrusts on facings in a GRS are smaller than in a GMSE

(Adams et al., 2011). These advantages offered by having closer-spaced reinforcements

are not accounted for in the SM. Instead, GRS is customarily designed using the

guidelines provided by Adams et al. (2011). One possible way of “harmonizing” the two

different design procedures is to analyze both GRS and GMSE as composite structures

whereby a composite strength of the reinforced soil is determined for use in the design

process.

Another use of composite strength in a GRS/GMSE wall is to determine the

relationship between the results of a columnar GRS/GMSE load test to those of the more

realistic plane strain condition (Nicks et al., 2013; Iwamoto et al., 2013) in the field as

discussed in Section 2.4.1. Full scale GRS load tests performed under plane strain

conditions (Bathurst and Benjamin, 1990; Pham, 2009; Wu et al., 2006) are more

expensive and more involved to set up and perform. As a result, many researchers have

resorted to testing a column of GRS instead (Adams, 1997; Adams et al., 2007; Adams et

al., 2002; Elton and Patawaran, 2005; Mitchell, 2002).

38

Performance Test Program

3.1 Test Configuration

A total of 14 GRS mini-piers were constructed and tested by the FHWA Turner-

Fairbank Highway Research Center (TFHRC) between fall 2011 and summer 2012.

Eight of those tests were chosen for analysis and the general characteristics of these tests,

designated as “TF”, are summarized in Table 3-1. Tests TF-1 through TF-5, and TF-8

will not be presented in this thesis; however their details can be found in “Geosynthetic

Reinforced Soil Performance Testing – Axial Load Deformation Relationships (Nicks et

al., 2013).

Highlighted herein are four pairs of load tests performed on square columnar

GRS/GMSE mini-piers. Each pair was identical in every sense except one test was

performed with cast masonry units (CMU) to serve as facing during load testing and the

other was performed with the CMU removed prior to load testing.

A schematic of each mini-pier pair (TF-6 and-7, TF-9 and -10, TF-11 and -12, and

TF-13 and -14) are shown in Figures 3-1 through 3-4. Tests TF-6, -9, -12, and -14 were

conducted with CMU blocks in place whereas TF-7, -10, -11 and -13 have similar

corresponding set-ups except the blocks were removed prior to testing. Plan schematics

shown in Figure 3-1 are the same for all mini-pier pairs. TF-14 had the same area in plan

but was slightly taller {H = 6.56 ft} with 7 pairs of full- and half-height blocks.

Photographs of a mini-pier with CMU, without CMU, and TF-14 are shown in Figure 3-5.

All test columns had a soil height to width ratio of ≈ 2.

39

Table 3-1 Mini-pier load tests conducted at TFHRC

Test Tf1

(lb/ft)

Sv2

(in)

Tf/Sv

(ksf)

Height of

Mini-Pier

(ft)

H/B Facing

Type

TF-6 4800 x 4800 7.63 7.55 6.35 1.95 CMU

TF-7 4800 x 4800 7.63 7.55 6.35 1.95 None

TF-9 4800 x 4800 15.3 3.78 6.35 1.95 CMU

TF-10 4800 x 4800 15.3 3.78 6.35 1.95 None

TF-11 1400 x 1400 3.81 4.41 6.35 1.95 None

TF-12 1400 x 1400 3.81 4.41 6.35 1.95 CMU

TF-13 3600 x 3600 11.3 3.84 6.56 2.02 None

TF-14 3600 x 3600 11.3 3.84 6.56 2.02 CMU

Notes: Tf = wide width tensile strength (all geotextiles have identical wide

width tensile strengths in the machine and cross machine directions)

Sv = reinforcement spacing

40

Figure 3-1 (a) Plan and profile schematic of TF-6; and (b) plan and profile schematic of

TF-7

a) b) POTS 1-4

Concrete Footing

Earth Pressure Cell

3 ft

4.56 ft

3.26 ft

Fatb

ack C

ell

Fatback Cell

6.3

5 ft

PO

TS

5

PO

TS

6

PO

TS

7

PO

TS

8

PO

TS

9

PO

TS

5-9

Concrete Footing

3 ft

3.26 ft

POTS 1-4

LV

DT

5

LV

DT

6

LV

DT

7

LV

DT

8

LV

DT

9

Earth Pressure Cell LV

DT

5-9

SG

1-1

S

G2-1

S

G3

-1

SG

1-2

S

G2-2

S

G3

-2

SG

1-3

S

G2

-3

S

G3

-3

3.26 ft

SG

1-1

S

G2

-1

S

G3

-1

SG

1-2

S

G2-2

S

G3-2

SG

1-3

S

G2

-3

S

G3

-3

SG1 SG2 SG3

SG4

SG5

SG1 SG2 SG3

SG4

SG5

41

Figure 3-2 (a) Schematic of TF-9; and (b) schematic of TF-10

a) b) POTS 1-4

Concrete Footing

Earth Pressure Cell

3 ft

6.3

5 ft

PO

TS

5

PO

TS

6

PO

TS

7

PO

TS

8

PO

TS

9

Concrete Footing

3 ft

3.26 ft

POTS 1-4

LVDT 5

LVDT 6

LVDT 7

LVDT 8

LVDT 9 Earth Pressure Cell

3.26 ft

Fatback Cell

42


a) b) POTS 1-4

Concrete Footing

Earth Pressure Cell

3 ft

4.56 ft

3.26 ft

Fatb

ack C

ell

Fatback Cell

6.3

5 ft

PO

TS

5

PO

TS

6

PO

TS

7

PO

TS

8

PO

TS

9

PO

TS

5-9

Concrete Footing

3 ft

3.26 ft

POTS 1-4

LV

DT

5

LV

DT

6

LV

DT

7

LV

DT

8

LV

DT

9

Earth Pressure Cell LV

DT

5-9

SG

1-1

S

G2-1

S

G3

-1

SG

1-2

S

G2-2

S

G3

-2

SG

1-3

S

G2

-3

S

G3

-3

3.26 ft

SG

1-1

S

G2

-1

S

G3-1

SG

1-2

S

G2-2

S

G3-2

SG

1-3

S

G2

-3

S

G3

-3

SG1 SG2 SG3

SG4

SG5

SG1 SG2 SG3

SG4

SG5

43


Figure 3-5 (a) Photograph of mini-pier test with CMU; (b) photograph of mini-pier test

without CMU; (c) photograph of TF-14

6.5

6 ft

a) POTS 1-4

Concrete Footing

3 ft

b) POTS 1-4

Concrete Footing

Fatback Cell

Earth Pressure Cell

3.26 ft

PO

TS

5

PO

TS

6

PO

TS

7

PO

TS

8

PO

TS

9

LV

DT

5

LV

DT

6

LV

DT

7

LV

DT

8

LV

DT

9

3.26 ft

Earth Pressure Cell

a) b) c)

44

3.2 Backfill

The backfill was a well-graded gravel with silty fines (GW-GM) from Lucky Stone

quarry in Leesburg, VA. The backfill met the requirements of Virginia Department of

Transportation’s (VDOT) 21A base course and FHWA specifications for use in GRS-IBS

abutments. The VDOT 21A gradation is shown in Figure 3-6.

Figure 3-6 VDOT 21A grain size distribution

Using a large scale shear box at FHWA's TFHRC, direct shear tests were run on

unscalped samples of this soil compacted at optimum based on Standard Proctor

{maximum dry unit weight = 147 pcf and optimum water content = 7.5%} at normal

stresses ranging from 5 psi to 30 psi in accordance with ASTM D3080. The compaction

curve is shown in Figure 3-7. The 12-inch x 12-inch x 8-inch high samples were sheared

at a rate of 0.015 in/min with a gap equal to D85 of the material (i.e. the aggregate size

where 85% of the sample is smaller). The Mohr-Coulomb failure envelope is shown in

Figure 3-8.

0

20

40

60

80

100

0.0010.010.11

Pe

rce

nt

Pa

ss

ing

(%

)

Grain Size (in.)

45

Figure 3-7 Compaction curve of VDOT 21A

Figure 3-8 Mohr-Coulomb failure envelope of GRS backfill based on large scale direct

shear tests

Assuming a linear Mohr-Coulomb (MC) envelope for the range of normal stresses

utilized, a friction angle and cohesion based on the measured peak strength during testing

were ≈ 54 and c ≈ 1560 psf, respectively for the partially saturated sample. When the

134

136

138

140

142

144

146

148

5% 6% 7% 8% 9% 10%

Dry

Un

it W

eig

ht

(pc

f)

Moisture Content (%)

Peak: = tan53° + 1563 psfR² = 0.9877

Fully Softened: = tan54°

R² = 0.9956

0

1000

2000

3000

4000

5000

6000

7000

8000

0 1000 2000 3000 4000 5000

Sh

ea

r S

tre

ss

(p

sf)

Normal Stress (psf)

Unsaturated - Peak MC EnvelopeSaturated - Peak MC EnvelopeUnsaturated - Fully Softened MC Envelope

46

samples were saturated prior to shear, the cohesion reduced to 125 psf. The MC

envelope is actually curved. A linear envelope was drawn to provide average shear

strength parameters of the backfill over the range of normal stresses utilized. Iwamoto et

al. (2013) showed that fully softened strengths of the backfill yielded capacities that

agreed better with measured GRS/GMSE capacities. A rationale for this is that the strong

reinforcement strengthens the soil considerably causing the GRS/GMSE to experience

large strains (> 10%) prior to failure whereas the peak strengths of the soils utilized in

these load tests were mobilized at relatively small shear strains (≈ 2-5%) in large scale

direct shear or triaxial tests. The fully softened was similar but c ≈ 0. This will be

discussed in more detail in the chapter entitled “Use of fully softened versus peak

strengths to predict the bearing capacity of a footing on GRS”.

Compaction was performed on each 3.8-inch-thick lift of the mini-pier using a

Vibco Patchman PM1012 plate compactor. The top two rows of blocks were ratchet

strapped together as compaction proceeded. Short pieces of 2 x 4 lumber were used to

increase contact between the straps and the blocks. Figure 3-9 shows the compaction of

the GRS mini-pier using this method. A nuclear gauge was employed to verify that the

relative compaction for each lift was close to 100% based on standard Proctor and that

the water content was within ±2% of optimum.

47

Figure 3-9 Ratchet straps and lumber used during compaction of GRS mini-pier

3.3 Geosynthetic Reinforcement

All geotextiles utilized were biaxial, woven polypropylene. However, they differed

in strength, stiffness and spacing as detailed in Table 3-2. The geotextiles were selected

with the aim of having a wide range of tensile strength to spacing ratio.

While the geotextiles used in each test were biaxial, the stiffnesses in the cross-

machine and machine directions of the reinforcement were different. For this reason, the

reinforcement was placed in an alternating pattern with each subsequent layer to prevent

preferential failure of the performance test in the weaker reinforcement direction.

48

Table 3-2 Properties of geotextiles used in the mini-pier load or performance tests

conducted at TFHRC.

Test Geotextile

Manufacturer

Wide Width

Tensile

Strength

(lb/ft)

Wide width

tensile strength

at 5% strain

(lb/ft)

Tensile

Strength

(Grab)

(lb)

Wide

Width

Elongation

(%)

TF-6, -7,

TF-9, -10 Propex 4800 x 4800 660 x 1500 600 x 500 (10 x 8%)

TF-11, -12 Industrial

Fabrics 1400 x 1400 Not specified 200 x 200 (9 x 7%)

TF-13, -14 US Fabrics 3600 x 3600 1392 x 1740 450 x 350 (15 x 10%)

3.4 Facing Elements

Each dry-cast and split-faced CMU block was 7.625 inches high x 15.625 inches

long x 7.625 inches wide (Figure 3-9 ) and weighed about 42 lbs on average. They were

frictionally connected to the geotextile reinforcement, with a coverage ratio of at least 85

percent as specified by Adams et al. (2011a). For mini-piers with CMU removed prior to

loading, the geotextiles were trimmed to be flush with the backfill. TF-13 and -14 used

alternating full- and half-height blocks to achieve a spacing of 11.3 inches. The top view

of the half block is identical. Its height is 3.8 inches.

Figure 3-9 Schematic of CMU block

3.5 Loading System

Load was applied by means of two 12-inch-stroke Enerpac jacks in series {giving a

total stroke of 24 inches} mounted on a two-post reaction frame that was bolted into a

7.63 in

15.63 in

Top view Side view

49

strong floor (Figure 3-10). A spherical bearing was placed on the footing centroid to

maintain a normal load on the concrete footing. Each load increment was manually

applied with the aid of a solenoid valve hydraulic pump. Load was maintained with a

strain indicator box calibrated to a load cell placed within the reaction assembly.

Hydraulic jack pressure was also checked at each load increment to ensure reliable

operation of the system throughout the course of a load test. At each increment, the load

was increased only when there was no significant change {< 0.003 inches} in settlement

between any two recording times as stated below; however, the load increment was held

for a minimum of 5 minutes and a maximum of 30 minutes. The data acquisition system

was programmed to record settlements at 1, 3, 5, 7, 15, 20, 25, and 30 minute intervals

from the start of each load increment. Typically, each test took about 6 hours to complete.

Figure 3-10 Loading system

Enerpac Jacks

Reaction frame

Data Acquisition

System

50

3.6 Instrumentation

The following is a list of instruments utilized during the load tests along with their

purposes.

Table 3-3 Instruments utilized in TF tests.

Instrument Purpose

Potentiometers and LVDT To measure lateral deflection of the GRS/GMSE

column along its height on one face

Potentiometers To measure vertical deflection of the footing at the top

of the GRS/GMSE column

Fatback Cell To measure lateral earth pressure at the CMU face near

the GRS/GMSE column mid-height

Earth Pressure Cell To measure vertical earth pressure near the bottom of

the GRS/GMSE column

Strain Gages To measure strains in two orthogonal directions in

three geotextile layers

3.6.1 Lateral and Vertical Deflection

Four string potentiometers (POTs), mounted on two reference beams, were

utilized to measure vertical settlement of the footing during load testing (Figure 3-11).

String potentiometers, mounted on a reference column, were utilized to measure lateral

displacements in the tests with CMUs. The potentiometers were replaced by linear

variable differential transducers (LVDTs) in the tests without CMUs (Figure 3-12). The

settlement of the footing and lateral displacement of one face of the mini-pier were

recorded every minute within each load increment. Movements were recorded during the

load test but not during construction of the test columns.

51

(a) (b)

Figure 3-11 Schematics of deflection instrumentation for mini-piers (a) with CMU and

(b) without CMU

3.6.2 Lateral and Vertical Earth Pressure

6-inch-diameter lateral earth pressure cells (Geokon Model 4810), also known as

Fatback cells, were located on the sixth block from the floor in tests TF-6, -9 and -12.

Figure 3-12 shows the installation of the Fatback cell onto the CMU block. In Test TF-

14, the Fatback cell was located on the fifth full block from the floor. Specifically

designed to measure soil pressures against a structural face, the Fatback cells have a thick

plate that stiffens the back of the cell so that its stiffness is more compatible with that of

the structure that it is mounted to. The performance tests were carried out to failure and

POTS 5

POTS 6

POTS 7

POTS 8

POTS 9

POTS 1-4

REFERENCE BEAMS

POTS 1-4

REFERENCE BEAMS

REFERENCE COLUMN

52

consequently, the Fatback cells were deemed not re-useable after each test. Due to

economic reasons, only one Fatback cell was used per test with facing.

Figure 3-12 Fatback cell mounted onto the CMU block.

Single-use 9-inch-diameter earth pressure cells (Geokon Model 4815) were also

placed horizontally at the centroid of the load test column in plan and at a fill height of

7.625 inches from the bottom to measure the vertical pressure in the soil at that elevation.

Vertical and lateral earth pressure measurements were collected at the end of each lift

during construction and at the end of each load increment.

3.6.3 Strain Gauges

Strain gauges were utilized to measure the strains in the geotextile during loading.

Five strain gauges were mounted on three of the geotextile layers in each mini-pier. The

strain gauges (type EP-08-250BG-120) were manufactured by Vishay Measurements

Group, Inc. and glued to a 1 inch x 3 inch rectangular piece of geotextile at two ends to

avoid stiffening of the geotextile if glued along its entire length. The rectangular patch

53

was then attached to the geotextile following the strain gauge attachment. This technique

was developed by the University of Colorado at Denver (Figure 3-13). The plan and

profile view of the strain gauge configuration is shown in the mini-pier schematics

(figures 3-1 through 3-4). The strain gauges were set up such that the strains in both

directions of the geotextile were measured.

Figure 3-13 Attachment of strain gauge on geotextile using the University of Colorado at

Denver attachment technique

54

Performance Test Results

Presented in this chapter are the results from four pairs of instrumented GRS/GMSE

mini-pier load tests with and without CMU blocks, designated as “TF”.

4.1 Ultimate Bearing Capacity

The ultimate bearing capacity, qult,emp, and the strain at failure, f are summarized in

Table 4-1. Figure 4-1 presents the load-settlement curves of the GRS mini-pier load tests.

In three pairs of tests, Tf/Sv = 4.095 0.315 ksf while the fourth pair had a Tf/Sv ratio =

7.55 ksf where Tf = reinforcing tensile strength and Sv = reinforcing spacing. The

following observations are offered:

Comparing the load tests with and without facing, the tests with CMU have higher

capacities due to the confinement provided by the CMU to the GRS.

The ultimate capacity increased significantly as the ratio Tf/Sv increased from 4.095

to 7.55 ksf (Figure 4-1). However, for a given value of Tf/Sv = 4.095 0.315 ksf, the

ultimate capacity varied amongst the three pairs of tests suggesting that the contribution

of the reinforcing tensile strength does not have the same weight as that of the fabric

spacing. In fact, it has been shown by Adams (1997), Adams et al (2007), Elton and

Patawaran (2004), Ziegler et al (2008) and Pham (2009) that spacing has a greater

influence on GRS capacity than tensile strength.

The ratio of the ultimate capacities with and without CMUs increased with

increasing reinforcement spacing and reinforcement tensile strength as shown in figures

4-2 and 4-3. This means that the CMU contribution to capacity increases with increasing

55

reinforcing spacing and reinforcing strength; i.e.; the smaller the reinforcing spacing and

the weaker the reinforcing, the smaller will be the effect of the CMU blocks on the

capacity.

TF-6 and -7 had a Tf and Sv similar to those used in the design of GRS-IBS bridge

abutments in the United States (Adams et al, 2011). TF-6 had the greatest ultimate

capacity, which was 11 times greater than the allowable bearing pressure of 4 ksf (Elias

and Christopher, 1997), and 65% greater than the corresponding unconfined test, TF-7.

TF-10 had the lowest ultimate capacity; however, it was still higher than the

allowable bearing pressure by a factor of 2.58. Note that the reinforcement spacing in

TF-9 and -10 is greater than the prescribed GRS maximum of12 in (Adams et al, 2011).

Table 4-1 Ultimate bearing capacity and strain at failure of mini-pier load tests

Test

Tf1

(lb/ft)

Sv2

(in)

Tf/Sv

(ksf)

Height

(ft)

Facing

Type

qult,emp3

(ksf)

qult,CMU4/

qult,No CMU5

f6

%

TF-6 4800 7.63 7.55 6.35 CMU 43.8 1.65

15.7

TF-7 4800 7.63 7.55 6.35 None 26.5 12.5

TF-9 4800 15.3 3.78 6.35 CMU 22.3 2.17

15.6

TF-10 4800 15.3 3.78 6.35 None 10.3 14.3

TF-11 1400 3.81 4.41 6.35 None 23.2 1.25

12.8

TF-12 1400 3.81 4.41 6.35 CMU 29.0 13.4

TF-13 3600 11.3 3.84 6.56 None 13.0 1.82

12.3

TF-14 3600 11.3 3.84 6.56 CMU 23.6 12.7

Notes:

1. Tf = wide width tensile strength (all geotextiles have identical wide width tensile

strengths in the machine and cross machine directions)

2. Sv = reinforcement spacing

3. qult,emp = ultimate bearing capacity of mini-pier

4. qult, CMU = ultimate bearing capacity of mini-pier with CMU

5. qult, No CMU = ultimate bearing capacity of mini-pier without CMU

6. εf = strain of load test at failure

56

Figure 4-1 Ultimate capacity versus Tf/Sv


spacing

43.8

22.3

29.0

23.6 26.6

10.3

23.3

12.9

10

15

20

25

30

35

40

45

0 1000 2000 3000 4000 5000 6000 7000 8000

Ult

ima

te C

ap

ac

ity (

ks

f)

Tf/Sv (psf)

CMU

No CMU

2.17

1.24

1.82

1.65

y = 0.6302ln(x) + 0.3813R² = 0.972

0.0

0.5

1.0

1.5

2.0

2.5

0 5 10 15 20

Ult

ima

te C

ap

ac

ity W

ith

C

MU

/Ult

ima

te C

ap

ac

ity W

ith

ou

t C

MU

Reinforcement Spacing, Sv (inches)

Tf/Sv = 4095psf

57


strength

The load-settlement curves for these tests are plotted in Figure 4-4. When the

applied vertical pressure is normalized by qult,emp and the settlement is normalized by the

settlement when q = qult,emp, the data plot within a fairly narrow range as shown in Figure

4-5.

2.17

1.24

1.82

1.65

y = 0.714ln(x) - 3.9476R² = 0.9767

0.0

0.5

1.0

1.5

2.0

2.5

0 1000 2000 3000 4000 5000 6000

Ult

ima

te C

ap

ac

ity W

ith

CM

U/U

ltim

ate

C

ap

ac

ity W

ith

ou

t C

MU

Reinforcement Strength, Tf (lb/ft)

Tf/Sv = 4095 psf

Tf/Sv = 7550 psf

58

Figure 4-4 Load-settlement curves of the mini-pier load tests

0

2

4

6

8

10

12

14

0 10 20 30 40 50S

ett

lem

en

t (i

n)

Applied Vertical Pressure (ksf)

TF-6 (Tᵳ/Sᵥ = 7.55 ksf, CMU) TF-7 (Tᵳ/Sᵥ = 7.55 ksf, No CMU)

TF-11 (Tᵳ/Sᵥ = 4.41 ksf, No CMU) TF-12 (Tᵳ/Sᵥ = 4.41 ksf, CMU)

TF-13 (Tᵳ/Sᵥ = 3.84 ksf, No CMU) TF-14 (Tᵳ/Sᵥ = 3.84 ksf, CMU)

TF-9 (Tᵳ/Sᵥ = 3.78 ksf, CMU) TF-10 (Tᵳ/Sᵥ = 3.78 ksf, No CMU)

59

Figure 4-5 Dimensionless form of the load-settlement curves of the mini-pier load tests

4.2 Failure Plane

In the tests without CMUs, distinct failure planes with ruptures in the geotextile

were observed when the reinforcement spacing was less than 12 inches (e.g.; figures 4-6

and 4-7).

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0.0 0.2 0.4 0.6 0.8 1.0S

ett

lem

en

t/S

ett

lem

en

t a

t q

ult

,em

p

q/qult,emp

TF-6 (Tᵳ/Sᵥ = 7.55 ksf, CMU)

TF-7 (Tᵳ/Sᵥ = 7.55 ksf, No CMU)







60


(Nicks et al, 2013)


(Nicks et al, 2013)

61

When the spacing exceeded 12 inches, no shear failure surface was evident in TF-

10 (Figure 4-8). According to Nicks et al (2013), failure of the soil between the

reinforcement occurred but the reinforcement themselves did not shear.

Figure 4-8 TF-10 at Failure (qult,emp = 10.33 ksf) with Sv = 15-1/4 in and Tf = 4800 lb/ft

(Nicks et al, 2013)

In the tests with CMU, shear failure development was not visible during the test.

However, upon removal of the CMU blocks at the end of the test, tear trends in the

geotextile provided some clues on the shear failure development. Pictures of individual

reinforcement layers are shown in Figure 4-9(a) through (i) for TF-6. Tears along all four

CMU edges occurred in the reinforcement layers from 2 to 6 CMU courses below the

footing. The reinforcement in the 7th CMU course below the footing had a tear along one

CMU edge while the reinforcements below that have no tears along the CMU. Also,

62

there were tears in the middle of the reinforcement layers from 1 to 7 CMU courses

below the footing. These middle tears progressively translate from one face of the test

column diagonally down towards the opposite face (Figure 4-9j). This is akin to the

classical shear surface observed in triaxial specimens. Tear trends in the other tests with

CMU closely resemble this pattern.

63

Figure 4-9 (a) – (i) Rupture pattern for geotextiles in TF-6; (j) Schematic of tears in

geotextiles in TF-6

a) Geotextile 1 c) Geotextile 3

d) Geotextile 4 e) Geotextile 5 f) Geotextile 6

g) Geotextile 7 h) Geotextile 8 i) Geotextile 9

b) Geotextile 2

Geotextile 9

Geotextile 8

Geotextile 5

Geotextile 6

Geotextile 7

Geotextile 4

Geotextile 3

Geotextile 1

Geotextile 2

j)

64

4.3 Lateral Pressures

4.3.1 During Mini-Pier Construction

Figure 4-10 presents measured lateral pressures as the fill height above the Fatback

cell increased during construction of mini-piers TF-6, -9, -12 and -14. In most tests, it

was observed that the lateral pressures during construction were small or negligible (< 0.2

ksf) and mostly less than the at-rest values for the unreinforced soil with ϕ = 54°. The

following observations are offered:

1. Compaction-induced stresses at the facing exceeded the soil at-rest values in TF-9

and TF-14 at shallow depths. However, as the depth increased, they dip below at-

rest values possibly due to (a) lateral displacement of the CMU blocks during

construction and/or (b) the geotextile restricts lateral movement of the soil, hence the

low lateral pressures on the CMU blocks. This implies that the reinforcement must

increase the soil’s apparent cohesion.

2. The fact that the lateral pressures are very small is consistent with the suggestions

of Wu (2001) and Soong and Koerner (1997) that the lateral stress on the facing of a

GRS/GMSE is proportional to the reinforcement spacing rather than a function of

the wall height and that for a constant reinforcement spacing, the lateral pressure

distribution is nearly uniform.

The highest lateral pressures were observed in TF-14, which had the second largest

spacing. However, TF-14 was also constructed during the warmest time of the year. It is

known that the 6-inch-diameter Fatback cells have a relatively high surface area to

volume ratio and thus may be temperature sensitive. The dates of construction of the

65

mini-pier and the temperatures recorded in the Fatback cells are summarized in Table 4-2.

The Fatback cell temperatures were more variable during construction of the mini-pier

than during load testing. Some lateral pressures were even negative which implies they

are essentially zero due to thermal effects. Since the temperatures remained fairly

constant during load testing and since the temperature-induced pressures are small

relative to the load-test-induced values by 1 to 2 orders of magnitude, temperature is thus

not expected to be a significant influencing factor in the lateral pressure trends during

testing as discussed below.

Table 4-2 Performance tests dates and Fatback cell temperatures

Test Date

Fatback Cell

Temperature

During Mini-Pier

Construction

(˚C)

Fatback Cell

Temperature

During Load

Testing

(˚C)

TF-6 December 2011 9.8-13.5 14.9-19.5

TF-9 February 2012 11.9-18.5 18.5-18.8

TF-12 April 2012 17.7-21.1 20.9-21.1

TF-14 June 2012 24.3-26.7 23.6-24.2

66

Figure 4-10 Measured lateral pressures at Fatback cell during construction of Mini-Piers.

4.3.2 During Load Testing

Measured lateral pressures on the CMU facing during load testing are presented in

Figure 4-11. As an example in TF-14, the measured horizontal stress increased with

increasing vertical load to about 1.63 ksf followed by a decrease. It is postulated that the

mechanism for the development of this confining stress is that as the footing is jacked,

the soil settles causing the geotextile to turn downward. This downturn of the fabric

exerts an axial load on the CMU blocks, which restrains the blocks from lateral

movement. Consequently, there is a build-up in lateral pressure giving rise to an

immensely strong and robust structure. Finally, as the load in the geotextile approaches

its tensile strength, the geotextile ruptures along the perimeter of the CMU blocks (Figure

0

5

10

15

20

25

30

35

40

-0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5F

ill H

eig

ht

Ab

ove

Fa

tba

ck

Ce

ll (

in)

Measured Lateral Pressure (ksf)

TF-6 (Tᵳ/Sᵥ = 7550 psf)

TF-12 (Tᵳ/Sᵥ = 4410 psf)

TF-14 (Tᵳ/Sᵥ = 3840 psf)

TF-9 (Tᵳ/Sᵥ = 3780 psf)

At-rest pressures with φ = 54˚

67

4-9). Its load is then shed to other reinforcement layers. A series of tears and load

spreads eventually causes the CMUs to lurch forward and the lateral pressure to reduce

abruptly. When the blocks are no longer part of the soil/geotextile/CMU composite, the

behavior tends towards a test without CMUs and the soil/fabric composite eventually

shears.

Figure 4-11 Measured lateral pressures at Fatback cell location during load testing of TF-

6, -9, -12, and -14

From Figure 4-11, the lateral stress is largest when Tf/Sv is highest. For a Tf/Sv =

7.55 ksf in TF- 6, up to 6.32 ksf of lateral stress was recorded before the Fatback cell

failed. In the other 3 tests where Tf/Sv ≈ one half that for TF-6, the lateral stresses never

-2

0

2

4

6

8

10

12

14

0 10 20 30 40 50

Me

as

ure

d L

ate

ral E

art

h P

res

su

re (

ks

f)

Applied Vertical Pressure (ksf)

TF-6 (Tᵳ/Sᵥ = 7550 psf)

TF-6 Extrapolated

TF-6 Kₐ

TF-12 (Tᵳ/Sᵥ = 4410 psf)

TF-12 Extrapolated

TF-12 Kₐ

TF-14 (Tᵳ/Sᵥ = 3840 psf)

TF-14 Kₐ

TF-9 (Tᵳ/Sᵥ = 3780 psf)

TF-9 Kₐ

68

exceeded 2.09 ksf. However, at approximately the same Tf/Sv = 4.095 0.315 ksf (tests

TF-9, -12 and -14), the larger the spacing, the larger the lateral stress against the face.

Also shown in Figure 4-11 are lines representing active pressures for the

GRS/GMSE composites. The lateral pressures during load testing at or near failure are

consistent with the Rankine active lateral earth pressure coefficients, Ka, the derivation of

which is discussed in the section on Lateral Earth Pressure Coefficients below. Also in

Figure 4-11, the lateral pressures for TF-6 and TF-12 near failure had to be extrapolated

because the Fatback Cells stopped working prematurely. The methodology for the

extrapolation is discussed in Section 4.6.

4.3.3 Lateral Earth Pressure Coefficients

Lateral earth pressure coefficients are plotted versus lateral movement in Figure

4-12. Assuming that the cohesion, c, (tabulated in Table 4-3 and derived in Section 4.6)

for the geo-composite is fully mobilized, the lateral earth pressure coefficient can be

calculated by solving for K in the following expression:

h = Kav – 2cKa½ (4-1)

where h = measured horizontal stress in the fatback cell and v = applied vertical stress

at the top. Some cohesion, whether it is due to soil suction, geosynthetic-induced-lateral-

soil restraint or a combination of both, has to be mobilized from the onset because the

GRS/GMSE can stand vertically and sustain load after the CMUs are removed (tests TF-

7, -10, -11 and -13). From Figure 4-12, it is observed that the lateral earth pressure

coefficients at failure are simply the Rankine active values based on the composite

friction angles in Table 4-3 (presented in Section 4.6.2). . This consistency justifies the

69

values of friction angles of the GRS composite derived in Section 4.6.2. Therefore, as

the friction angle increases, Ka [= tan2(45˚ - /2)] decreases.

Figure 4-12 Lateral earth pressure coefficients versus dimensionless lateral movement

4.4 Lateral Deformation

Lateral deformations along the face of GRS columns were instrumented using

POTS and LVDTs in tests with and without facing, respectively. LVDTs were removed

before failure of TF-10 and -13 due to raveling of the soil around the LVDT reaction

plates. Reliable readings for these two tests were available only up to 83% of the

ultimate load. For consistency, lateral deformation readings for all tests are presented at

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

0.00 0.01 0.02 0.03 0.04

La

tera

l E

art

h P

res

su

re C

oe

ffic

ien

t

Lateral Deformation/Height (in/in)

Rankine Kₐ for TF-6Rankine Kₐ for TF-12Rankine Kₐ for TF-14Rankine Kₐ for TF-9TF-6 (Tᵳ/Sᵥ = 7550 psf)TF-12 (Tᵳ/Sᵥ = 4410 psf)TF-14 (Tᵳ/Sᵥ = 3840 psf)TF-9 (Tᵳ/Sᵥ = 3780 psf)TF-6 ExtrapolatedTF-12 Extrapolated

TF-12 Ka = tan2(45° - 30°/2) = 0.33

TF-6 Ka = tan2(45° - 11°/2) = 0.67

TF-9 Ka = tan2(45° - 32°/2) = 0.305

TF-14 Ka = tan2(45° - 33°/2) = 0.29

70

an applied load level of q approximately equal to 83% of the ultimate load, qult,emp in

Figure 4-13. Actual q/qult,emp values are labeled in Figure 4-13.

It can be observed that:

1. The maximum lateral deformation increased with decreasing spacing.

2. The maximum lateral deformation tends to occur in the middle of the pier

(sometimes a little higher) when the spacing were less than 12 inches. At a

reinforcement spacing of 15.3 inches, the maximum lateral deformation occurred

at or near the top. Directly below the footing where the load is highest, the soil-

reinforcement interface friction restrains lateral deformation of the soil. It also

results in higher locked-in lateral stresses in the GRS (not to be confused with

lateral stress on the CMU) after compaction. This explains why the deformation

is higher when there is less reinforcement at the top. There is an upward shift in

the location of maximum lateral deformation from tests without CMU to those

with CMU.

71

Figure 4-13 Lateral deformation profiles of GRS at loads ≈ 83% of ultimate load

Figures 4-14 and 4-15 show the lateral displacement along the height of the GRS

column at several loads throughout testing of TF-6 and -7. It is evident that:

1. Larger lateral strains are attained in the test with CMU.

2. The lateral movement is generally largest at mid-height; i.e.; the GRS face bellies

out. Initially, the entire deflected profile translates horizontally. After the

deflections at the top reach a maximum, they then reduce towards the end of the

test, and the deflected profile at larger loads cross over those at lower loads near

the top. This trend is true in both tests regardless of whether there is a facing or

not.

0

10

20

30

40

50

60

70

0 1 2 3 4D

ep

th b

elo

w t

op

of

wa

ll (

in)

Horizontal Displacment (in)

TF-6 (q/qᵤ = 0.82, Tᵳ/Sᵥ = 7.55 ksf, CMU)

TF-7 (q/qᵤ = 0.80, Tᵳ/Sᵥ = 7.55 ksf, No CMU)







72

Figure 4-14 Lateral displacement of TF-6 with increasing applied load (with CMU) Tf/Sv

= 7.55 ksf

Figure 4-15 Lateral deformation of TF-7 with increasing applied load (without CMU)

Tf/Sv = 7.55 ksf

0

10

20

30

40

50

60

70

-0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0D

ista

nc

e f

rom

to

p o

f w

all

(in

)Horizontal Displacment (in) .9 ksf

1.8 ksf2.5 ksf3.7 ksf4.5 ksf5.4 ksf6.3 ksf9.9 ksf13.4 ksf18.7 ksf21.4 ksf24.0 ksf26.7 ksf29.3 ksf30.6 ksf31.9 ksf33.3 ksf35.8 ksf38.7 ksf41.3 ksf43.8 ksf

0

10

20

30

40

50

60

70

-0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0

Dep

th b

elo

w t

op

of

wa

ll (

in)

Horizontal Displacement (in)

.9 ksf

1.8 ksf

2.7 ksf

3.6 ksf

4.4 ksf

5.4 ksf

6.2 ksf

8.0 ksf

9.8 ksf

10.7 ksf

13.4 ksf

16.0 ksf

18.7 ksf

21.3 ksf

24.1 ksf

26.6 ksf

73

4.4.1 Postulate of Zero Volume Change

Adams, et al. (2002) performed load tests on large scale GRS mini-piers (3.67ft x

3.67ft x 8ft high; Tf = 2400 lb/ft; Sv = 6 inches with 3 inch spacing in the top 4 layers;

segmental retaining wall block facing). Results from vertical and lateral deformation

data suggested a zero net volume change during vertical loading. Based on this and other

available tests from the literature, Adams et al (2002) postulated the theory of zero

volume change (ZVC). They suggested that this postulate can be used to predict the

maximum lateral deflection of a GRS structure. The validity of the postulate of zero

volume change is investigated in this section.

Adams, et al. (2002), calculated lateral expansion using Vlateral = 4 x Vface = 4 x

(Ho x W x DL/2), where Vface is the lateral expansion of the instrumented face of the mini

pier, Ho is the original height, W is the width, and DL is the maximum measured lateral

deformation. The lateral expansion along all faces of the mini-pier was assumed to be

equal. The vertical compression was calculated as Vtop = As x S, where As was the plan

area of the GRS and S was the average settlement of the GRS.

The lateral expansion of the TF tests was calculated differently. First, the height

of the GRS used to calculate the lateral expansion at each load increment was a function

of the measured vertical settlement. As an example, in Figure 4-16a, the lateral

deformation measured by LVDT-5 was not included in the calculation of the lateral

expansion because the GRS mass settled past the LVDT location at this load. Second,

the lateral deformation along the face of the GRS was integrated over the GRS height

which allowed for a more accurate representation of the lateral deformation of the GRS

74

with load increase. Finally, it was assumed that the corners of the GRS did not move or

expand with vertical compression, Figure 4-16b. Although expansion of the corners was

observed in photographic evidence of the tests, the movements were slight and not large

enough to discredit this assumption. All four faces of the GRS were assumed to deform

equally. Figures 4-16a and b show the difference in the assumed shape of the deformed

GRS mass (e.g. TF-11 at q = 23 ksf) between Adams, et al. (2002) and the TF tests.

Figure 4-16 Schematic of assumed deformed mass for TF-11 at applied load = 23 ksf

(drawn to scale) as assumed by Adams, et al. (2002) and TF tests in (a) profile view and

(b) plan view

q = 23 ksf

Original dimensions Ho = 76.25 in. W = 39.06 in.

Settlement ρ = 9.72 in

Adams, et al. (2002) Deformed GRS mass

(blue)

TF

Deformed GRS mass

(red)

DLmax

DLmax

1.98 in

LVDT-6 DL = 1.98 in

LVDT-7 DL = 1.88 in

LVDT-8 DL = 1.37 in

LVDT-9 DL = 0.52 in

DL

a) profile b) plan

75

Volumetric strain, εv, defined as volume change/ original volume, is plotted versus

vertical applied load in Figure 4-17. In this figure, it can be seen that:

1. The theory of zero volume change is true at service loads (less than 4 ksf) where

there is less than 0.7% change in volume in all tests (Figure 4-18). At applied

stresses larger than 4 ksf, the volume change does not exceed 5%.

2. The smaller the spacing (TF-11 and -12 have the smallest while TF-9 and -10

have the largest), the closer the behavior tends towards zero volume change.

3. All tests except TF-6 dilated with initial increase in vertical load. TF-14 dilated

the most (εv = 4.3%). TF-6 first compressed until about 10 ksf, then “dilated”

until about 36 ksf and then compressed again until failure.

4. Among the tests with CMU, the maximum dilation occurred at the maximum

measured lateral stress in TF-9. The GRS then experienced vertical compression

with additional increase in load, ultimately approaching zero volume change (v =

0). Among the tests with CMU, the maximum measured lateral stress occurred

just past the peak dilation in TF-14. Therefore, the peak dilation and the peak

lateral stress roughly coincided in these two tests. In TF-14, the fabric that

ruptured first probably did not coincide elevation-wise with the POT that recorded

the largest lateral deflection.

5. Among the tests with CMU, the Fatbacks failed prematurely in TF-6 and -12.

6. TF-9 and -14, and TF10 and -13 have similar deformation shapes with load

increase. These are the piers with the largest spacing with Sv = 11.3 and 15.3

inches in TF-13 and -14, and TF-9 and 10, respectively.

76

7. In the tests without CMU (TF-7, -10, -11, and -13), the GRS dilated and never

compressed when the spacing are large. However as the spacing decreased, (TF-7

and -11), the GRS deforms closer to the ZVC line.

Figure 4-17 Volumetric strain versus vertical applied load

Figure 4-18 presents the volumetric strain varying with applied load up to a

typical service load of 4 ksf. As indicated previously, |εv| < 0.7%. In this plot,

volumetric strains are smallest in tests without CMU. Also among these tests without

CMU, volumetric strain decreased with increasing Tf/Sv ratio. TF-6 only compressed and

-0.05

-0.04

-0.03

-0.02

-0.01

0

0.01

0.02

0.03

0.04

0.05

0 5 10 15 20 25 30 35 40 45

v

Vertical Applied Load (ksf)









Late

ral E

xp

an

sio

n

Ve

rtica

l Co

mp

ressio

n

Maximum

measured h

77

never expanded in this range of stress. The remaining tests all dilated in this range of

stress.

Figure 4-18 Volumetric strain up to the service load limit (4 ksf)

Figure 4-19 presents the volume gained (lateral strain) versus volume lost

(vertical strain) as a percent of the initial volume (lateral expansion versus vertical

compression). From this plot, it can be observed that without CMU, the “strain-paths”

“concaved” away from the ZVC line whereas the “strain-paths” of the tests with CMU

“convexed” toward the ZVC line. In other words, the CMU “contained” the GRS when

the GRS was subjected to large vertical loads.

-0.01

-0.005

0

0.005

0.01

0 1 2 3 4

v

Vertical Applied Load (ksf)









Late

ral E

xpansio

n

Ve

rtica

l Co

mp

ressio

n

78

Figure 4-19 Volume gained versus volume lost

4.5 Fabric Strains

4.5.1 Introduction

Strain measurements are useful for determining reinforcement loads in the fabric

when subjected to loading. However, appropriate modulus of the geotextile necessary for

converting strains to loads must be obtained under the in-service confining stress rather

than in air, where data is most widely available. The modulus can be measured at

varying confining stress with the aid of an air bag plus some nominal soil cover.

However, this data is not readily available. Therefore, the objective of this section is to

0

0.05

0.1

0.15

0.2

0 0.05 0.1 0.15 0.2

Vg

ain

ed/V

o

VLost/Vo









79

observe trends in the measured fabric strains in the TF tests, and to comment on some of

the results.

4.5.2 Strain Gauge Layout

Strains in three levels of geotextile reinforcing in both the machine and cross-

machine directions were measured in the performance tests. Table 4-3 summarizes the

strain gage locations. Schematics of the instrumented geotextiles in the GRS piers are

shown in figures 3-1 through 3-4.

Table 4-3 Summary of strain gages in performance tests

Test Sv

(in)

Instrumented

Geotextile1

Instrumented Geotextile

Between CMU Blocks2

TF-6 and -7 7.6

3

5

7

3 and 4

5 and 6

7 and 8

TF-9 and -10 15.3

2

3

4

4 and 5

6 and 7

8 and 9

TF-11 and 12 3.8

6

10

14

3 and 4

5 and 6

7 and 8

TF-13 and -14 11.3

2

3

5

2nd ½ block and 3rd full block

3rd ½ block and 4th full block

5th ½ block and 6th full block Notes: 1. Geotextile numbering counted from bottom up. This numbering

system is used in this section.

2. CMU block numbering counted from bottom up.

As an example, the numbering of the strain gages is as follows in TF-6 and -7.

The third strain gauge of an instrumented geotextile sheet, say geotextile layer L, or SGL-

3 was located at the centroid of the square geotextile; SGL-4 and -5 extended 7.5 and 15

inches, respectively out towards the face in the cross-machine direction. SGL-2 and -1

extended 7.5 and 15 inches, respectively, out toward the face in the machine direction.

80

The machine direction of the geotextile also coincides with the direction of lateral

deformation measurements.

4.5.3 Results

Strain measurements were available during the early stages of loading with most

gauges failing once the applied pressure exceeded ~ 4 ksf. These strains are plotted

against lateral distance from the center of the GRS (or SGL-3) in Figures 4-20(a) – (c)

through 4-22(a) – (c) for TF-6, -11 and -12, respectively. Results for tests TF-7, -9, -10, -

13 and -14 were sparser and have been omitted. Also plotted on the far left are the lateral

strains estimated using lateral deformations at the facing measured using POTs (tests with

CMU) or LVDT (tests without CMU). These lateral strains at the facing can be

compared to strains from the strain gages. The lateral strain at the facing can be

calculated as the ratio of the lateral deformation to half the width of the GRS.

The following observations are offered:

TF-6 (test with the largest Tf/Sv ratio and with CMU):

Fabric strains increased with increasing applied stress in all tests. Fabric strains

were larger in the machine direction than the cross-machine direction suggesting

that there was eccentricity in the loading.

Strains calculated from the POTs were less than the fabric strains indicating that

some of the geotextile strain energy is used to overcome friction between the

fabric and the CMU and between the fabric and the soil.

81

Geotextile strains are largest in the middle of the test column than at the top. This

is also true in tests TF-9 and -11 but not -12.

TF-11 (test with the smallest reinforcement spacing and without CMU):

Fabric strains are symmetric in the machine and cross-machine directions when

the applied stresses reached 4.37 ksf or 19% of the failure capacity (23.2 ksf).

This suggests that this column is very uniformly loaded up to this point.

The calculated strains from the LVDTs are lower than the fabric strains,

suggesting that a large portion of the strain energy goes into overcoming friction

between the soil and fabric only, since this test was conducted without CMU.

TF-12 (test with the smallest reinforcement spacing and with CMU):

Fabric strain distribution is also fairly symmetric in the machine and cross-

machine directions, as observed in TF-11.

Strains were highest at the center of the GRS for top and middle geotextile layers.

It could be inferred that no shear plane has developed yet at this stage of the test.

The strain magnitude in this test is lower than in all the other tests. Due to the

small reinforcement spacing, not much load has transferred down to the geotextile

at an applied pressure of 3.08 ksf.

From the observations discussed above, it can be deduced that:

1. The type of strain gages used does not last very long into the load test. They

generally fail at an applied pressure of approximately 4 ksf. The longer they last,

the more useful they can be for indicating the direction of the failure plane.

82

2. The strain gages provide a good indication of the concentricity of the applied load

and may be useful for real-time load adjustments.

3. In many of the tests, the largest strains occur close to the test column mid-height.

This agrees with the fact that the GRS column tends to deflect or barrel most at

mid-height.

4. For very closely spaced reinforcement (~4 inches), the strain distribution between

the machine and cross-machine directions are fairly uniform at small load levels.

A uniform strain distribution in a geotextile is an indication that development of a

shear plane has not begun.

83


center of the GRS mass for (a) Geotextile 7 (b) Geotextile 5 and (c) Geotextile 3 for TF-6

POT6

SG3-1SG3-2

SG3-3SG3-4

SG3-5

-0.2

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)

Distance from center (in)

0.89 ksf

1.76 ksf

2.52 ksf

3.72 ksf

POT7

SG2-1SG2-2

SG2-3

SG2-4

SG2-5

-0.2

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)

Distance from Center (in)

0.89 ksf1.76 ksf2.52 ksf3.72 ksf

POT8 SG1-1SG1-2

SG1-3SG1-4

SG1-5

-0.2

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)


0.89 ksf1.76 ksf2.52 ksf3.72 ksf

84


center of the GRS mass for (a) Geotextile 14 (b) Geotextile 10 and (c) Geotextile 6 for

TF-11

LVDT6

SG3-1SG3-2

SG3-3

SG3-4

SG3-5

0.0

1.0

2.0

3.0

4.0

5.0

-15-7.507.51522.530

Str

ain

(%

)


1.32 ksf1.71 ksf2.68 ksf3.51 ksf4.37 ksf

a)

LVDT7

SG2-1SG2-2

SG2-3

SG2-4 SG2-5

0.0

1.0

2.0

3.0

4.0

5.0

-15-7.507.51522.530

Str

ain

(%

)


1.32 ksf

1.71 ksf

2.68 ksf

3.51 ksf

4.37 ksf

b)

LVDT8

SG1-1SG1-2

SG1-3SG1-4

0.0

1.0

2.0

3.0

4.0

5.0

-15-7.507.51522.530

Str

ain

(%

)


1.32 ksf

1.71 ksf

2.68 ksf

3.51 ksf

4.37 ksf

c)

85


center of the GRS mass for (a) Geotextile 16 (b) Geotextile 12 and (c) Geotextile 8 for

TF-12

POT6SG3-1

SG3-2

SG3-3

SG3-4SG3-5

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)


0.99 ksf

2.02 ksf

3.08 ksf

a)

POT7

SG2-1SG2-2

SG2-3

SG2-4SG2-5

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)


0.99 ksf

2.02 ksf

3.08 ksf

b)

POT8

SG1-3 SG1-4SG1-5

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-15-7.507.51522.530

Str

ain

(%

)


0.99 ksf

2.02 ksf

3.08 ksf

c)

86

4.6 Shear Strength Parameters of a GRS/GMSE

For a given backfill, it can be shown that the shear strength parameters of a

GRS/GMSE composite compacted to the same physical state varies with Tf and Sv. The

load test results can be used to discern shear strength parameters of pairs of

corresponding load tests with and without CMUs by constructing their stress paths.

Stress path of the MIT (Lambe and Whitman, 1969) variety {p = ½(1 + 3) and q =

½(1 - 3)} is used herein but there is no reason why other types cannot be utilized. The

MIT stress path represents the top of the Mohr circle during the progression of a load test.

Without CMU, there is no applied lateral pressure or confining stress and the stress path

rises from the origin along a 45˚ line to the right till failure as shown in Figure 4-23 for

TF-13. Also shown in Figure 4-23 is the stress path for TF-14 which is identical in all

respects to TF-13 except that CMUs provide confinement. The stress paths for the other

three pairs of tests with and without CMUs are shown in figures 4-24 (TF-9 and -10), 4-

25 (TF-11 and -12) and 4-26 (TF-6 and -7). The procedure for plotting the stress path for

all tests with CMUs is described below.

87


CMU) with Tf/Sv = 3.84 ksf



Maximum σh

measured in Fatback Cell

Kf Lineq = tan31°p' + 3 ksf

0

2

4

6

8

10

12

14

16

18

0 5 10 15 20 25 30

, q

(k

sf)

σ, p' (ksf)

TF-14TF-14 Post-failureTF-13

TF-13 and 14 Kᵳ-LineTF-13 and 14 MC Envelope21A Peak MC Envelope21A Fully Softened MC Envelope

Maximum σh

measured in Fatback Cell

Kf Lineq = tan31°p' + 2.5 ksf

0

2

4

6

8

10

12

14

16

18

0 5 10 15 20 25 30

, q

(ks

f)

σ, p' (ksf)

TF-9TF-9 Post-failureTF-10TF-9 and 10 Kᵳ-LineTF-9 and 10 MC Envelope21A Peak MC Envelope21A Fully Softened MC Envelope

88





Best-fit curveR² = 0.999

Kf Lineq = tan29°p' + 6 ksf

0

2

4

6

8

10

12

14

16

18

0 5 10 15 20 25 30

, q

(k

sf)

σ, p' (ksf)

TF-12

TF-12 Extrapolated

TF-11

TF-11 and 12 Kᵳ-LineTF-11 and 12 MC Envelope

21A Peak MC Envelope

21A Fully Softened MC Envelope

Best-fit CurveR² = 0.999

Kf Lineq = tan11.5°p' + 10.7 ksf

0

2

4

6

8

10

12

14

16

18

0 5 10 15 20 25 30

, q

(k

sf)

σ, p' (ksf)

TF-6TF-6 ExtrapolatedTF-7TF-6 and 7 Kᵳ-LineTF-6 and -7 MC Envelope21A Fully Softened MC Envelope21A Peak MC Envelope

89

4.6.1 Construction of Stress Paths

One necessary assumption needed to plot the stress paths for tests with CMU is

that the lateral stress in the Fatback cell is representative of the confining stress along the

entire height of the GRS/GMSE. This assumption is made in light of the following (a)

Figure 4-10 shows that the lateral pressures are relatively small even when the height of

backfill increases; (b) Wu (2001) and Soong and Koerner (1997) suggest that for a

constant reinforcement spacing, the lateral pressure distribution is approximately constant

with depth; and (c) Mitchell (2002) showed that the lateral pressures measured in

performance tests were fairly equally distributed over the wall height for geotextile

spacings of 16 inches or less. Upon making this assumption, the entire GRS/GMSE

column can be idealized as a single element as in the case of a triaxial soil sample with

constant horizontal and vertical stresses, except that the vertical and horizontal stresses

are not principal stresses because vertical shear stresses act along the CMU and

complementary horizontal shear stresses act below the footing. The shear stresses can be

estimated as follows:

Like conventional retaining walls, "shear stresses" do develop along the soil-wall

interface of the mini-piers with CMUs. This shear force, T, embodies both the upward

wall friction between the soil and the facing and the upward tensile force in the geotextile,

the mechanism for which has been explained in Section 4.3.2. It can be estimated by

vertical force equilibrium (Figure 4-27) and by using the measured vertical stress from

the earth pressure cell near the bottom as follows:

90

T = Qapp + Wf + Ws - VEP

(4-2)

where Qapp = applied jacking load from the load cell, Wf = weight of footing, Ws =

weight of soil above the bottom earth pressure cell and VEP = upward vertical reaction

force = vertical stress measured by the bottom earth pressure cell x area of soil in the

GRS/GMSE. In a separate free body diagram (Figure 4-27), it can be shown that the

vertical downward components of wall friction and of load in the geotextile, plus the

weight of the CMU blocks are resisted by the upward reaction of the CMUs with the

strong floor. Hence, they do not factor into Equation 3. The shear stress, , can then be

calculated as:

= T/ACMU

(4-3)

where ACMU = vertical area of the CMU wall above the earth pressure cell. It is also

assumed that is representative of the shear stress at the wall mid-height.

At each load test increment, the Mohr circle passes through (h, ) and (v, )

where h = measured horizontal stress in the Fatback cell, v = applied vertical stress at

the top and = shear stress at the Fatback cell elevation calculated using Equation 4-3. In

all cases, it was found that was quite small; i.e.; v and h ≈ major and minor principal

stresses, respectively.

91

Figure 4-27 Free-body diagram of vertically loaded GRS

The lateral pressures in the Fatback reach a peak and then decrease when the

geotextiles rupture along the facing periphery as seen in Figure 4-11 for TF-9 and -14.

At this juncture, the CMU blocks lurch forward and offer little confinement to the geo-

composite. The behavior then tends towards a test without CMU as there is then a shift

in direction of the stress path towards the 45˚ line as seen in figures 4-23 and 4-24. If

failure is defined as the point when the horizontal stress is a maximum, the Kf-line which

Geotextile 9

Geotextile 8

Geotextile 5

Geotextile 6

Geotextile 7

Geotextile 4

Geotextile 3

Geotextile 1

Geotextile 2

T

Qapp

Wf

Ws

VEP

Rfloor

Wblock

Rfloor

Wblock

T

T = Qapp + Wf + Ws − VEP

T = Rfloor − Wblock

92

is the equivalent of the Mohr-Coulomb envelope in p-q space, can be drawn by joining

this point to the last point of the corresponding test without CMU; i.e.; it is assumed that

the failure envelope is linear and the shear strength parameters for the geo-composite is

the same for the load tests with and without CMUs. Essentially, the CMUs merely serve

as a means to apply a variable confining stress to the GRS/GMSE.

Using the slope (tan ) and intercept (m) of the Kf-line, the corresponding shear

strength parameters for the GRS/GMSE can be computed as follows (Lambe and

Whitman, 1969):

= sin-1(tan ) (4-4)

c = m/cos 4-5)

4.6.2 Strength Parameters

The composite shear strength parameters are summarized in Table 4-3. It should

be noted that the Fatback cells in tests TF-6 and -12 stopped working prior to failure and

extrapolation of the stress path using a second order polynomial was necessary to discern

the shear strength parameters for these tests. However, the extrapolated results should

not be too erroneous based on the fact that: (1) the vertical stress is known throughout the

test and it constrains the values of p and q; and (2) photographic evidence of when the

CMU blocks lurch forward is available to “identify” the failure load step, examples of

which are shown in figures 4-28 (a) and (b) for TF-6 and (c) and (d) for TF-12.

93

Table 4-4 Shear strength parameters of GRS

Test Tf

(lb/ft)

Sv

(in)

Tf/Sv

(ksf)

Facing

Type

Composite Shear

Strength Parameters

c

(ksf)

ϕ

(°)

TF-6 4800 7.63 7.55 CMU 10.9 11

TF-7 4800 7.63 7.55 None

TF-9 4800 15.3 3.78 CMU 2.9 32

TF-10 4800 15.3 3.78 None

TF-11 1400 3.81 4.41 None 6.7 30

TF-12 1400 3.81 4.41 CMU

TF-13 3600 11.3 3.84 None 3.6 33

TF-14 3600 11.3 3.84 CMU

Figure 4-28 (a) and (b) Photographic evidence of when CMU blocks lurched forward at

failure of TF-6; (c) and (d) Photographic evidence of when CMU blocks lurched forward

at failure of TF-12

c) 26.2 ksf

a) 41.25 ksf b) 43.8 ksf – Failure Load

d) 29.1 ksf – Failure Load

94

4.6.3 Discussion on the Shear Strength Parameters

Dimensionless forms of cohesion and friction angle are plotted versus a

dimensionless form of the Tf/Sv ratio, Tf/(Svpa) where pa = atmospheric pressure in Figure

4-29. From Figures 4-23 through 4-26, Figure 4-29, and Table 4-3, the following

observations can be made:

The Mohr-Coulomb envelopes for the GRS/GMSE are not parallel to those for the

unreinforced soil, nor are they parallel to each other (unless they have similar Tf/Sv

ratios) despite having similar soil unit weights and water contents. This is contradictory

to Yang’s (1972) and Schlosser and Long’s (1974) assertion that the friction angles of the

unreinforced and reinforced soil are identical.

The friction angle of the GRS/GMSE composite is less than that of the

unreinforced soil. In general, the composite friction angle increases with decreasing Tf/Sv

(decreasing Tf and increasing Sv).

The cohesion of the GRS/GMSE is considerably higher than that of the

unreinforced soil. It increases with decreasing spacing and increasing geotextile strength.

The benefits of reinforcing a soil become significant if Sv is small since the smaller the Sv,

the further the failure envelope extends above that for the unreinforced soil (Figure 4-23

through 4-26).

95

Figure 4-29 Dimensionless forms of cohesion and friction angle versus dimensionless for

of Tf/Sv ratio

0.1

0.2

0.3

0.4

0.5

0.6

0.7

2

4

6

8

10

12

14

1.5 2.0 2.5 3.0 3.5 4.0

tan

G

RSc

GR

S/

H

Tf/SvPa

96

Use of Fully Softened Versus Peak Strengths to Predict the Ultimate

Bearing Capacity of Footings on GRS

5.1 Peak versus Fully Softened Strengths

A database of 28 GRS load tests were collected from the literature and from this

study to investigate whether the fully softened or the peak shear strength is more

appropriate for use in predicting the ultimate carrying capacity of footings on GRS. The

work in this chapter has been published by Iwamoto et al (2013 and 2014). Table 5-1

summarizes details of these tests, most of which were discussed in sections 2.1, 3 and 4.

Zornberg (2002) proposed that peak strengths of the backfill be used to design GRS

slopes based on centrifuge model testing. Christopher et al. (1990) proposed that peak

strengths of the backfill be used to predict the bearing capacity of footings on

mechanically stabilized earth (MSE) type bridge abutments reinforced using extensible

elements. For the soils used to construct the GRS in this database, it was observed that

the peak strength from large scale direct shear (LSDS) tests (Figure 5-1a) on granular

soils was mobilized at about 0.51 to 0.63 inch lateral displacements, which correspond to

about 4.5 to 5.4% shear strain for a 12 in x 12 in x 8 in high direct shear sample.

Additionally, the peak strengths from large scale (6 inch diameter x 12 inch high) triaxial

tests on granular soils (Figure 5-1b) was mobilized at about 2.3 to 5.0% axial strain

(Pham, 2009). In this study, it was observed that load tests on GRS with closely spaced

(< 12 inch) reinforcement having a wide width tensile strength of at least 4800 lb/ft

generally fail at strains greater than 10%, implying that the soil’s shear strength then is

past its peak value.

97

Table 5-1 Test parameters of GRS performance tests selected from literature

Test

dmax

in

USCS

Symbol

Strength

Test Type

Tf

lb/ft

Sv

in

Facing

Type

h

psi

Boundary

Conditions

GSGC21 1.3 GS-GM TX2 4800 7.63 3 4.9 PS4

GSGC31 1.3 GS-GM TX2 9600 15.25 3 4.9 PS4

GSGC41 1.3 GS-GM TX2 4800 15.25 3 4.9 PS4

GSGC51 1.3 GS-GM TX2 4800 7.63 None 0.0 PS4

Elton1 0.5 SP DS5 612 6.00 None 0.0 CC6







VS-17 0.5 GP LSDS8 2400 7.63 CMU 4.4 SC9

VS-27 0.5 GP LSDS8 4800 7.63 CMU 4.4 SC9

VS-57 0.5 GP LSDS8 4800 7.63 CMU 4.4 SC9

MPA11 0.8 GW-GM LSDS8 4800 15.25 None 4.4 SC9

MPB11 0.5 GW-GM LSDS8 4800 7.63 None 4.4 SC9

MPC11 1.0 GW-GM LSDS8 4800 7.63 None 0.0 SC9

TF-1 1.0 GP LSDS8 4800 7.63 CMU 0.0 SC9

TF-2 1.0 GW-GM LSDS8 1400 7.63 CMU 0.0 SC9

TF-3 0.5 GW-GM LSDS8 2400 7.63 None 4.4 SC9








TF-14 1.0 GW-GM LSDS8 2400 11.44 CMU 4.4 SC9 Notes: 1. GSGC = Generic Soil-Geosynthetic Composite

2. TX = Consolidated drained triaxial compression tests on 0.15-m-diameter and 0.3-m-high

samples

3. No facing was used. Instead, a confining pressure = 5 psi was applied using a rubber

membrane wrapped around the GRS

4. PS = Plane strain boundary conditions

5. DS = direct shear test. Because soil was SP, direct shear sample was 2.5 in. diameter

performed in accordance with ASTM D3080

6. CC = Cylindrical Column

7. VS = Performance tests conducted in Defiance County, OH as part of the FHWA’s Every

Day Counts GRS Validation Sessions

8. LSDS = Large Scale Direct Shear tests on 12 in by 12 in by 8 in high specimen

9. SC = Square Column

10. MP = Mini pier tests or more widely referred to herein as performance tests

98

a) b)

Figure 5-1 (a) Large scale direct shear device (LSDS) (b) Large scale triaxial device

(LSTX)

Figure 5-2 Typical stress-strain curve of an overconsolidated soil

The notion of using fully softened strength is not new in geotechnical engineering.

Duncan and Wright (2005) recommend the use of fully softened strength when analyzing

the stability of cuts in heavily overconsolidated soil. The rationale for this is that

swelling and softening was found to have occurred along the slip surfaces during forensic

studies of such slides and use of fully softened strength, or the strength if the soil was

normally consolidated, provided better agreement when back-calculating the factors of

safety in these failed slopes. Wu (1996) indicated that the fully softened strength is

typically mobilized at strains on the order of 10% (Figure 5-2). Clearly soil swelling is

irrelevant here. In the case of GRS, the reinforcement strengthens the soil and forces

Peak strength

Fully softened strength

Residual strength

ε ≈ 2 - 5% ε ≥ 10%

𝜎𝑑

99

failure to occur very often at double-digit strains (Table 5-2). Therefore, in the interest of

preserving strain compatibility, a study was conducted to see whether fully softened

strengths provide a better prediction of the bearing capacity of footings on GRS

abutments than peak strengths.

The ultimate bearing capacities of the 28 load tests were predicted using Equation

2-3 and both the backfill’s peak and fully softened shear strength parameters (as shown in

Table 5-2). Figure 3-8 shows a plot of the fully softened and peak Mohr-Coulomb

envelopes for the Virginia 21A well-graded soil used in this study.

5.1.1 Use of Peak Strengths to Predict Bearing Capacity

The predicted ultimate bearing capacities using the soil’s peak, qult,peak, and fully

softened, qult,fs, strengths and the measured ultimate capacities (qult,emp) of the 28 load

tests are shown in Table 5-2. Also shown is the bias defined as the ratio of the measured

to predicted ultimate capacities. The mean, standard deviation, and coefficient of

variation (COV) of the biases are shown at the bottom of Table 5-2.

The average bias, defined as the measured capacity divided by the predicted

capacity, calculated using peak strengths was 0.79 (COV = 36%). Figure 5-4 shows the

histogram and the probability density function (PDF) of the normal distribution of the

bias. Figure 5-5 shows the predicted versus measured ultimate capacities using peak

strengths. Clearly, Equation 2-3 over-predicts the GRS capacity when using peak

strengths.

100

Table 5-2 Predicted and measured ultimate bearing capacity of GRS load tests using fully

softened versus peak strengths

Test

qult,emp

psf

εf1

% ϕpeak

˚

cpeak

psi

qult,peak

psf

ϕfs

˚

cfs

psi

qult,fs

psf

Bias, λ

Peak FS

GSGC2 70957 6.5 50 10 51299 41 17 46429 1.35 1.80

GSGC3 42574 6.1 50 10 42183 41 17 37438 1.04 1.33

GSGC4 37252 4 50 10 31265 41 17 26669 1.38 1.62

GSGC5 42449 6 50 10 47174 41 17 42360 0.90 1.18

Elton1 4805 1.7 40 4 5351 41 4 5266 0.92 0.91

Elton2 2695 3.1 40 4 3215 41 4 3125 0.86 0.86

Elton3 6392 3.9 40 4 6959 41 4 6879 0.94 0.93

Elton4 6100 4.5 40 4 7126 41 4 7045 0.88 0.87

Elton5 8398 4.7 40 4 8401 41 4 8324 1.02 1.01

Elton6 8293 7.7 40 4 8901 41 4 8825 0.95 0.94

Elton7 9589 8.5 40 4 10398 41 4 10326 0.94 0.93

VS-1 23310 8 54 3 38891 51 0 28829 0.70 0.96

VS-2 22709 7.1 46 3 31419 45 0 27537 0.82 0.94

VS-5 21539 10.4 51 0 28829 51 0 28829 0.89 0.89

MPA 4696 1.9 54 11 14989 53 0 5559 0.31 0.87

MPB 3548 2.2 54 11 22676 53 0 13316 0.16 0.27

MPC 9600 6.4 54 11 22363 53 0 13000 0.43 0.76

TF-1 20487 10.9 53 10 27834 55 0 21101 0.88 1.34

TF-2 25260 11.5 54 112 37331 53 0 28105 0.79 1.15

TF-3 17491 13.8 54 112 31565 53 0 22286 0.55 0.81

TF-6 43763 15.7 54 112 59416 53 0 50391 0.81 1.00

TF-7 26546 12.5 54 112 53650 53 0 44572 0.49 0.61

TF-9 22310 15.6 54 112 29282 53 0 19982 0.94 1.60

TF-10 10330 14.3 54 112 23516 53 0 14164 0.44 0.75

TF-11 23249 12.8 54 112 41800 53 0 32615 0.56 0.73

TF-12 29030 13.4 54 112 47566 53 0 38433 0.69 0.90

TF-13 12960 12.3 54 112 27086 53 0 17767 0.48 0.75

TF-14 23562 12.7 54 112 32853 53 0 23585 0.85 1.32

Average Bias 0.79 1.00

Standard Deviation of the Bias 0.28 0.32

Coefficient of Variation (%) 0.36 0.32 Notes: 1. εf = Strain of GRS load tests at failure

2. Best fit linear Mohr-Coulomb envelopes for soil used in TF-2 through TF-14

yielded values of cohesion of 11 psi and 1 psi for the partially saturated and

saturated samples, respectively. A cohesion of 11 psi was used when

estimating the GRS capacity since the soil was partially saturated during load

testing in FHWA’s TFHRC laboratory.

101

Figure 5-3 Histogram and normal distribution of the bias using peak strengths

Figure 5-4 Measured versus predicted capacities using peak strengths

0

0.3

0.6

0.9

1.2

1.5

0

1

2

3

4

5

0 0.5 1 1.5 2

Pro

ba

bility

Den

sity

Fu

nctio

n

Fre

qu

en

cy

l (qult,emp/qult,peak)

0

10

20

30

40

50

60

70

80

0 20 40 60 80

qu

lt,p

eak

(ks

f)

qult,emp (ksf)

GSGC (Pham, 2009)

Elton (Elton and Patawaran, 2005)

VS (Nicks et al., 2013)

MP (Adams et al., 2007)

TF (Nicks et al., 2013)

102

5.1.2 Use of Fully Softened Strengths to Predict Bearing Capacity

In contrast, the average bias of the measured capacity to the predicted capacity

using fully softened strengths was 1.00 (COV = 32%). Figure 5-6 shows the histogram

and PDF of the normal distribution of the bias; Figure 5-7 shows the predicted versus

measured capacities using fully softened strengths. In Figure 5-7, the data points are

more centered on the line of equality. Based on these results, Equation 2-3 with the use

of fully softened strengths yield a bias that is close to unity with a slightly smaller COV

compared to the use of peak strengths for this dataset.

Hypothesis testing on the normal distribution of the bias indicate that using

Equation 2-3 and fully softened strengths will result in a mean bias of 1.00, with a 90%

confidence that the bias will be within 3 standard deviations of the mean.

Figure 5-5 Histogram and normal distribution of the bias using fully softened strengths

0

0.3

0.6

0.9

1.2

1.5

0

1

2

3

4

5

0 0.5 1 1.5 2

Pro

ba

bility

Den

sity

Fu

nctio

n

Fre

qu

en

cy

l (qult,emp/qult,fs)

103

Figure 5-6 Measured versus predicted capacities using fully softened strength

5.2 Summary

Fully softened strengths are more suitable for bearing capacity predictions because

GRS with closely spaced reinforcement generally fail at large strains past the backfill’s

peak strengths. A follow-on to this is that since large movements are required to fail, say

a GRS abutment, the design of GRS abutments will most likely be governed by the

serviceability limit state rather than the ultimate limit state

0

10

20

30

40

50

60

70

80

0 20 40 60 80

qu

lt,f

s(k

sf)

qult,emp (ksf)

GSGC (Pham, 2009)

Elton (Elton and Patawaran, 2005)

VS (Nicks et al., 2013)

MP (Adams et al., 2007)

TF (Nicks et al., 2013)

104

Summary and Conclusions

6.1 Summary

Four pairs (one with and one without CMU facing) of GRS load tests, also known

as mini-piers and denoted as “TF” in this study, were tested at the TFHRC. Tests were

instrumented to measure the vertical load, vertical and lateral deformation, and vertical

earth pressure near the bottom and lateral earth pressures at mid-height of the mini-piers

with CMU and geotextile fabric strains. The mini-pier pairs had varying Tf/Sv ratio (TF-

6 and -7 Tf/Sv = 7.55 ksf; TF-9 and -10 Tf/Sv = 3.78 ksf; TF-11 and -12 Tf/Sv = 4.41 ksf;

TF-13 and -14 Tf/Sv = 3.84 ksf). This study focused on observations of the vertical

capacity of GRS columns in particular the effects of CMU facing on capacity and also

their load-settlement curves, lateral earth pressures during mini-pier construction and

during load testing to failure, the validity of the postulate of zero volume change (Adams

et al., 2002), composite shear strength parameters of GRS, and the appropriateness of

using the soil’s fully softened versus peak shear strengths to predict the bearing capacity

of footings on GRS.

6.2 Findings and Conclusions

The following findings and conclusions from this study are offered:

6.2.1 Bearing Capacity

Comparing the load tests with and without facing, the tests with CMU have higher

capacities due to the confinement provided by the CMU to the GRS.

The ultimate capacity increased significantly as the ratio Tf/Sv increased.

However, for a given value of Tf/Sv, the ultimate capacity varied amongst three

105

pairs of tests having different combinations of Tf and Sv suggesting that the

contribution of the reinforcing tensile strength does not have the same weight as

that of the fabric spacing. In fact, it has been shown by other researchers that

spacing has a greater influence on GRS capacity than tensile strength.

The CMU contribution to capacity increased with increasing reinforcing spacing

and reinforcing strength; i.e.; the smaller the reinforcing spacing and the weaker

the reinforcing, the smaller will be the effect of the CMU blocks on the capacity

6.2.2 Failure Plane:

In the tests without CMUs, distinct failure planes with ruptures in the geotextile

were observed when the reinforcement spacings were less than 12 inches. Shear

failure development was not visible during the tests with CMU. However, upon

removal of the CMU blocks at the end of the test, tear trends in the geotextile

provided some clues on the shear failure development.

When the spacing exceeded 12 inches, no shear failure surface was evident.

Failure of the soil between the reinforcement occurred but the reinforcement

themselves did not shear.

6.2.3 Lateral Pressures:

For the CMU facing utilized, measured lateral pressures were very small or

negligible and less than the at-rest values of the unreinforced soil during GRS

construction.

106

Measured lateral pressures increased during load testing and are largest when

Tf/Sv is highest. However, at approximately the same Tf/Sv, the larger the spacing,

the larger the lateral stress.

A behavioral mechanism for the load tests is theorized as follows. As the footing

is loaded, the soil settles causing the geotextile to turn downward. This downturn

in the geotextile exerts an axial force on the CMU blocks, which restrains the

blocks from lateral movement. As a result, there is a build-up in lateral pressure

giving rise to an immensely strong and robust structure. Finally, as the load in the

geotextile approaches its tensile strength, the geotextile ruptures along the

perimeter of the CMU blocks. When this happens, the lateral pressure in the

CMU drops. The CMU blocks are now no longer part of the geo-composite.

Instead the behavior now tends towards the test without any CMU blocks. With a

loss in confining stress, the soil/fabric composite then shears more readily.

6.2.4 Lateral Deformation:

The maximum lateral deformation increased with increasing spacing.

The maximum lateral deformation tends to occur in the middle of the pier

(sometimes a little higher) when the spacings were less than 12 inches. At a

reinforcement spacing greater than 12 inches, the maximum lateral deformation

occurred at or near the top. Soil-reinforcement interface friction restrains lateral

deformation of the soil. It also results in higher locked-in lateral stresses in the

GRS. This explains why the top deformation is higher when there is less

reinforcement at the top.

107

6.2.5 Postulate of Zero Volume Change

The theory of zero volume change is true at service loads (less than 4 ksf) where

there is less than 0.7% change in volume in all tests. At applied stresses larger

than 4 ksf, the volume change does not exceed 5%.

The smaller the reinforcement spacing, the closer the behavior tends towards zero

volume change.

6.2.6 Fabric Strains

The type of strain gages used does not last very long into the load test. They

generally fail at an applied pressure of approximately 4 ksf. The longer they last,

the more useful they can be for indicating the direction of the failure plane.

The strain gages provide a good indication of the concentricity of the applied load

and may be useful for real-time load adjustments.

In many of the tests, the largest strains occur close to the test column mid-height.

This agrees with the fact that the GRS column tends to deflect or barrel most at

mid-height.

For very closely spaced reinforcement (~4 inches), the strain distribution between

the machine and cross-machine directions are fairly uniform at small load levels.

A uniform strain distribution in a geotextile is an indication that development of a

shear plane has not begun.

108

6.2.7 Shear Strength Parameters

The Mohr-Coulomb envelopes for the GRS/GMSE composites are not parallel to

those of the unreinforced soil nor are they parallel to each other despite having

similar soil densities and water contents.

The friction angle of the GRS/GMSE composite is less than that of the

unreinforced soil. In general, the composite friction angle increases with

decreasing Tf/Sv (decreasing Tf and increasing Sv). The same factors affecting

friction angle also affect the derived lateral earth pressure coefficients in the same

manner.

Geosynthetic reinforcement introduces a significant cohesion to the soil that is

higher than that of the unreinforced soil. The cohesion increases with decreasing

spacing and increasing geotextile strength.

The benefits of reinforcing a soil become increasingly significant when Sv

decreases since the smaller the Sv, the further the failure envelope extends above

that for the unreinforced soil.

6.2.8 Fully Softened versus Peak Strengths

Fully softened strengths are more suitable for bearing capacity predictions

because GRS with closely spaced reinforcement generally fail at large strains past

the backfill’s peak strengths.

A follow-on to this is that since large movements are required to fail say a GRS

abutment, the design of GRS abutments will most likely be governed by the

serviceability limit state rather than the ultimate limit state

109

6.3 Key Findings

Key contributions from this study include: (1) the derivation of the soil-

geosynthetic composite shear strength parameters and (2) the use of the soil’s fully

softened strength to predict the bearing capacity of a footing on a GRS is more

appropriate than the peak.

Use of stress paths and measured lateral and vertical earth pressures to deduce the

composite shear strength parameters of a GRS has not been performed previously. From

this study, it was observed that the soil-geosynthetic composite friction angle is less than

that of the unreinforced soil, contrary to what has been assumed by several researchers

(Schlosser and Long, 1972; Yang, 1972). The composite friction angle decreases with

increasing Tf/Sv ratio and increases with decreasing Tf and increasing Sv. The

geosynthetic reinforcement also introduces a significant cohesion to the soil that is higher

than that of the unreinforced soil. The cohesion increases with decreasing Sv and

increasing Tf.

Use of fully softened strengths improves the accuracy and reliability of Pham’s

(2009) equation used for predicting the ultimate bearing capacity of a footing on a GRS.

6.4 Recommendations for Future Works

1. Lateral earth pressures were measured using one Fatback cell installed mid-height

and mid-width of one face of the GRS. The lateral pressure at that one location

was assumed to be equal along the height of the wall and for all faces. Installing

more Fatback cells along all faces of the wall would allow for verification of this

assumption.

110

2. This study considered only the measured vertical and lateral stresses at failure

using only two GRS/GMSE tests and assumed a linear failure envelope when

deriving the composite shear strength parameters. It does not consider the

kinematic effects of the load tests. Conceivably more rationally sound composite

shear strength parameters can be back calculated using numerical analyses to

match both the measured vertical and lateral deflections and stresses.

3. The TF test series were built as square columns whereas a footing on a GRS

abutment more resembles a plane strain condition. The bearing capacity equation

using GRS composite shear strength parameters was deemed to be applicable.

However, the effects of the corners formed by the CMU of the square columns on

the prediction of GRS behavior when subjected to vertical load should be

investigated. This may be performed using numerical analyses such as the finite

element method, which is beyond the scope of this thesis. The load tests reported

herein were conducted using only one soil type, essentially one wall height, one

type of CMU block and woven polypropylene geotextiles. To fully develop a

prediction method for the composite shear strength of a GRS/GMSE, additional

tests involving different soils, wall heights, reinforcement and facing types may

be of interest.

111

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Documents

OBSERVATIONS FROM LOAD TESTS ON GEOSYNTHETIC REINFORCED SOIL