Geosynthetic Reinforced Soil 

Performance Testing—

Axial Load Deformation Relationships

Publication No. FHWA-HRT-13-066 AUGUST 2013

Research, Development, and Technology

Turner-Fairbank Highway Research Center

6300 Georgetown Pike

McLean, VA 22101-2296
FOREWORD

The use of geosynthetic reinforced soil (GRS) for load bearing applications such as bridge 

abutments and integrated bridge systems (IBS) has expanded among transportation agencies 

looking to save time and money while delivering a better and safe product to the traveling public. 

GRS has been identified by the Federal Highway Administration (FHWA) as a proven, marketready technology, and is being actively promoted through its Every Day Counts (EDC) initiative. 

FHWA interim design guidance for GRS abutments and IBSs is presented in Publication No. 

FHWA-HRT-11-026. The guidance includes the procedure and use of the GRS performance 

tests, also termed a mini-pier experiment. This report presents a database of nineteen 

performance tests performed by the FHWA, largely at the Turner-Fairbank Highway Research 

Center. It also presents findings, conclusions, and suggestions regarding various design 

parameters related to the performance of GRS, such as backfill material, reinforcement strength, 

reinforcement spacing, facing confinement, secondary reinforcement, and compaction.

A reliability analysis for load and resistance factor design (LRFD) was performed based on the 

results of this performance testing to determine a calibrated resistance factor for the soilgeosynthetic capacity equation. The results of this analysis can also be used by bridge designers 

to estimate capacity and deformation of GRS. In addition, an insight into the behavior of GRS as 

a new composite material due to the close reinforcement spacing is described.

Jorge E. Pagán-Ortiz

Director, Office of Infrastructure

Research and Development

Notice

This document is disseminated under the sponsorship of the U.S. Department of Transportation 

in the interest of information exchange. The U.S. Government assumes no liability for the use 

of the information contained in this document. This report does not constitute a standard, 

specification, or regulation.

The U.S. Government does not endorse products or manufacturers. Trademarks or 

manufacturers’ names appear in this report only because they are considered essential to the 

objective of the document.

Quality Assurance Statement

The Federal Highway Administration (FHWA) provides high-quality information to serve 

Government, industry, and the public in a manner that promotes public understanding. Standards 

and policies are used to ensure and maximize the quality, objectivity, utility, and integrity of its 

information. FHWA periodically reviews quality issues and adjusts its programs and processes to 

ensure continuous quality improvement.
TECHNICAL REPORT DOCUMENTATION PAGE

1. Report No.

FHWA-HRT-13-066

2. Government Accession No. 3. Recipient’s Catalog No.

4. Title and Subtitle

Geosynthetic Reinforced Soil Performance Testing—Axial Load 

Deformation Relationships

5. Report Date

August 2013

6. Performing Organization Code: 

HRDI-40

7. Author(s)

Nicks, J.E., Adams, M.T., Ooi, P.S.K., Stabile, T.

8. Performing Organization Report No. 

9. Performing Organization Name and Address

Turner-Fairbank Highway Research Center

6300 Georgetown Pike

McLean, VA 22101

10. Work Unit No.

11. Contract or Grant No.

N/A

12. Sponsoring Agency Name and Address

Office of Infrastructure R&D

FHWA Research, Development and Technology

6300 Georgetown Pike

McLean, VA 22101

13. Type of Report and Period Covered

Technical

14. Sponsoring Agency Code

15. Supplementary Notes

The FHWA Contracting Officer’s Technical Representative (COTR) was Mike Adams, HRDI-40.

16. Abstract

The geosynthetic reinforced soil (GRS) performance test (PT), also called a mini-pier experiment, consists of 

constructing alternating layers of compacted granular fill and geosynthetic reinforcement with a facing element 

that is frictionally connected, then axially loading the GRS mass while measuring deformation to monitor 

performance. This large element load test provides material strength properties of a particular GRS composite 

built with unique combinations of reinforcement, compacted fill, and facing elements. This report describes the 

procedure and provides axial load- deformation results for a series of PTs conducted in both Defiance County, 

OH, as part of the Federal Highway Administration’s (FHWA) Every Day Counts (EDC) GRS Validation 

Sessions and in McLean, VA, at the FHWA’s Turner-Fairbank Highway Research Center as part of a parametric 

study.

The primary objectives of this research report are to: (1) build a database of GRS material properties that can be 

used by designers for GRS abutments and integrated bridge systems; (2) evaluate the relationship between 

reinforcement strength and spacing; (3) quantify the contribution of the frictionally connected facing elements at 

the service limit and strength limit states; (4) assess the new internal stability design method proposed by Adams 

et al. 2011 for GRS; and (5) perform a reliability analysis of the proposed soil-geosynthetic capacity equation for 

LRFD calibration.

(1,11)

17. Key Words

Geosynthetic reinforced soil, performance test, mini-pier 

experiment, abutment, integrated bridge system, geotextile, 

capacity, deformation

18. Distribution Statement

19. Security Classif. (of this report)

Unclassified

20. Security Classif. (of this page)

Unclassified

21. No. of Pages

169

22. Price

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized
ii

SI* (MODERN METRIC) CONVERSION FACTORS 

APPROXIMATE CONVERSIONS TO SI UNITS

Symbol When You Know Multiply By To Find Symbol 

LENGTH

in inches 25.4 millimeters mm 

ft feet 0.305 meters m 

yd yards 0.914 meters m 

mi miles 1.61 kilometers km 

AREA 

in2square inches 645.2 square millimeters mm2

ft2 square feet 0.093 square meters m

2

yd2 square yard 0.836 square meters m

2

ac acres 0.405 hectares ha

mi2square miles 2.59 square kilometers km2

VOLUME 

fl oz fluid ounces 29.57 milliliters mL 

gal gallons 3.785 liters L 

ft3 cubic feet 0.028 cubic meters m

3 

yd3 cubic yards 0.765 cubic meters m

3 

NOTE: volumes greater than 1000 L shall be shown in m3

MASS 

oz ounces 28.35 grams g

lb pounds 0.454 kilograms kg

T short tons (2000 lb) 0.907 megagrams (or "metric ton") Mg (or "t") 

TEMPERATURE (exact degrees) 

o

F Fahrenheit 5 (F-32)/9 Celsius oC 

or (F-32)/1.8 

ILLUMINATION 

fc foot-candles 10.76 lux lx

fl foot-Lamberts 3.426 candela/m2 cd/m2

FORCE and PRESSURE or STRESS 

lbf poundforce 4.45 newtons N 

lbf/in2poundforce per square inch 6.89 kilopascals kPa 

APPROXIMATE CONVERSIONS FROM SI UNITS 

Symbol When You Know Multiply By To Find Symbol 

LENGTH

mm millimeters 0.039 inches in 

m meters 3.28 feet ft 

m meters 1.09 yards yd 

km kilometers 0.621 miles mi 

AREA 

mm2square millimeters 0.0016 square inches in2 

m

2

square meters 10.764 square feet ft2 

m

2

square meters 1.195 square yards yd2 

ha hectares 2.47 acres ac

km2 square kilometers 0.386 square miles mi2 

VOLUME 

mL milliliters 0.034 fluid ounces fl oz 

L liters 0.264 gallons gal 

m

3 cubic meters 35.314 cubic feet ft3 

m

3 cubic meters 1.307 cubic yards yd3 

MASS 

g grams 0.035 ounces oz

kg kilograms 2.202 pounds lb

Mg (or "t") megagrams (or "metric ton") 1.103 short tons (2000 lb) T 

TEMPERATURE (exact degrees)

oC Celsius 1.8C+32 Fahrenheit o

F 

ILLUMINATION 

lx lux 0.0929 foot-candles fc 

cd/m2candela/m

2

0.2919 foot-Lamberts fl

FORCE and PRESSURE or STRESS 

N newtons 0.225 poundforce lbf 

kPa kilopascals 0.145 poundforce per square inch lbf/in2

*SI is the symbol for the International System of Units. Appropriate rounding should be made to comply with Section 4 of ASTM E380. 

(Revised March 2003)
iii

TABLE OF CONTENTS

1. INTRODUCTION.......................................................................................................................1

1.1 BACKGROUND ..............................................................................................................1

1.2 CURRENT PERFORMANCE TESTS ........................................................................4

1.3 INTERNAL STABILITY DESIGN ..............................................................................5

1.4 OBJECTIVES ..................................................................................................................9

2. TESTING CONDITIONS........................................................................................................11

2.1 BACKFILL CONDITIONS .........................................................................................11

2.1.1 Sieve Analysis.....................................................................................................12

2.1.2 Density ................................................................................................................15

2.1.3 Friction Angle .....................................................................................................15

2.2 REINFORCEMENT CONDITIONS .........................................................................17

2.3 FACING CONDITIONS ..............................................................................................17

3. TEST SETUP .............................................................................................................................19

3.1 LAYOUT.........................................................................................................................19

3.2 LOAD AND REACTION FRAME.............................................................................20

3.3 CONSTRUCTION.........................................................................................................25

3.4 INSTRUMENTATION.................................................................................................27

3.5 LOAD SCHEDULE AND COLLECTION OF DATA............................................28

4. RESULTS DATABASE............................................................................................................29

4.1 UNLOAD/RELOAD BEHAVIOR ..............................................................................33

4.2 REPEATABILITY ........................................................................................................38

4.3 COMPOSITE BEHAVIOR..........................................................................................38

5. COMPARISON TO PLANE STRAIN CONDITIONS ......................................................45

5.1 CAPACITY.....................................................................................................................45

5.2 STIFFNESS.....................................................................................................................48

6. PARAMETRIC ANALYSIS....................................................................................................53

6.1 EFFECT OF AGGREGATE TYPE ...........................................................................53

6.2 EFFECT OF COMPACTION .....................................................................................54

6.3 EFFECT OF BEARING BED REINFORCEMENT...............................................58

6.4 EFFECT OF GRADATION.........................................................................................61

6.5 EFFECT OF REINFORCEMENT STRENGTH.....................................................64

6.6 EFFECT OF THE RELATIONSHIP BETWEEN REINFORCEMENT 

STRENGTH AND SPACING................................................................................................65

6.7 EFFECT OF FACING..................................................................................................71

7. APPLICATIONS OF PERFORMANCE TESTING TO DESIGN ..................................79

7.1 DESIGN DATABASE...................................................................................................79

7.2 STRENGTH LIMIT......................................................................................................80

7.2.1 Analytical ............................................................................................................80

7.2.2 Empirical .............................................................................................................84

7.3 SERVICE LIMIT...........................................................................................................87
iv

7.4 LRFD CALIBRATION FOR STRENGTH LIMIT.................................................93

7.4.1 Background .........................................................................................................93

7.4.2 Reliability Analysis: FOSM...............................................................................94

7.4.3 Resistance Factor................................................................................................97

8. CONCLUSIONS........................................................................................................................99

APPENDIX A. SOIL TESTING DATA ....................................................................................103

APPENDIX B. NUCLEAR DENSITY TESTING FOR TFHRC PTS .................................115

APPENDIX C. DEFORMATION INSTRUMENTATION LAYOUTS FOR PTS ............125

APPENDIX D. RAW DATA FOR PTS .....................................................................................131

ACKNOWLEGEMENTS .............................................................................................................149

REFERENCES................................................................................................................................151
v

LIST OF FIGURES

Figure 1. Photo. Vegas mini-pier experiment.....................................................................................1

Figure 2. Illustration. Plan view of Vegas mini-pier experiment ......................................................2

Figure 3. Illustration. Face view of Vegas mini-pier experiment......................................................2

Figure 4. Illustration. Side view of Vegas mini-pier experiment ......................................................3

Figure 5. Illustration. Reinforcement schedule for Vegas mini-pier experiment .............................4

Figure 6. Illustration. Plan view of Defiance County experiment .....................................................5

Figure 7. Illustration. Elevation view of Defiance County (DC) test................................................5

Figure 8. Equation. Ultimate vertical capacity for a GRS composite ...............................................6

Figure 9. Equation. W factor................................................................................................................6

Figure 10. Equation. Required reinforcement strength ......................................................................6

Figure 11. Equation. Confining stress (Wu et al. 2010).....................................................................7

Figure 12. Graph. Predictive capability of the soil-geosynthetic composite capacity equation .....8

Figure 13. Graph. Predictive capability of the required reinforcement strength equation...............8

Figure 14. Graph. Reinforced backfill gradations............................................................................14

Figure 15. Graph. LSDS testing results.............................................................................................16

Figure 16. Photo. DC-1 GRS PT (before testing).............................................................................20

Figure 17. Illustration. Concrete footing on GRS composite, inset from facing............................21

Figure 18. Photo. Hollow core hydraulic jacks for PT assembly ....................................................21

Figure 19. Photo. TF-1 PT setup with reaction frame......................................................................22

Figure 20. Photo. TF-6 PT setup with reaction frame......................................................................23

Figure 21. Photo. TF-10 PT setup with reaction frame....................................................................23

Figure 22. Photo. TF-9 at failure with reaction frame......................................................................24

Figure 23. Photo. Spherical bearing to apply load to the footing on the GRS composite .............25

Figure 24. Illustration. Instrumentation layout for DC tests and TF-1............................................27

Figure 25. Illustration. General additional instrumentation layout TF PT series ...........................28

Figure 26. Graph. Load-deformation behavior for the Defiance County PTs................................30

Figure 27. Graph. Load-deformation behavior for the Turner Fairbank PTs.................................31

Figure 28. Photo. Tilting of the footing during TF-4 testing ...........................................................32

Figure 29. Graph. TF-4 results...........................................................................................................33

Figure 30. Graph. TF-1 results...........................................................................................................35

Figure 31. Graph. TF-5 results...........................................................................................................36

Figure 32. Graph. TF-6 results...........................................................................................................37

Figure 33. Graph. Repeatability of PT at TFHRC............................................................................39

Figure 34. Photo. TF-11 at failure with Sv = 3-13/16 inches, Tf = 1,400 lb/ft, and 21A 

material .............................................................................................................................40

Figure 35. Photo. TF-3 at failure with Sv = 75⁄8 inches, Tf = 2,400 lb/ft, and 21A material ..........40

Figure 36. Photo. TF-13 at failure with Sv = 111⁄4 inches, Tf = 3,600 lb/ft, and 21A material ......41

Figure 37. Photo. TF-10 at failure with Sv = 151⁄4 inches, Tf = 4,800 lb/ft, and 21A material. .....41

Figure 38. Photo. Rupture pattern for geotextiles in TF-6 (qult,emp = 43,828 psf); the lowest 

layer of reinforcement is the closet fabric in the picture ...............................................43

Figure 39. Photo. Post-test picture of TF-6 (Sv = 75⁄8 inches, Tf = 4,800 lb/ft)...............................44

Figure 40. Equation. Mohr-Coulomb shear strength........................................................................45

Figure 41. Equation. Ultimate capacity of an unconfined GRS PT ................................................45

Figure 42. Equation. Ultimate capacity of a strip footing on slope.................................................45
vi

Figure 43. Equation. Ultimate capacity of a strip footing on a vertical GRS abutment.................45

Figure 44. Equation. Ratio of plane strain capacity to PT capacity ................................................46

Figure 45. Equation. Stability Factor ................................................................................................46

Figure 46. Graph. Ratio of plane strain capacity to PT capacity for different stability factors ....46

Figure 47. Graph. Mohr-Coulomb failure envelope for Pham (2009) plane strain GSGC tests ..47

Figure 48. Graph. Plane strain capacity to PT capacity for a stability factor of 0.29 ....................48

Figure 49. Illustration. Infinitely Long Unconfined GRS abutment ...............................................48

Figure 50. Equation. Stiffness of an Infinitely Long Unconfined GRS abutment .........................49

Figure 51. Illustration. Solution for strip footing on top of a wall ..................................................49

Figure 52. Equation. Vertical displacement of a GRS abutment with a strip footing ....................50

Figure 53. Equation. Vertical strain...................................................................................................50

Figure 54. Equation. Stiffness of a GRS abutment supporting a strip footing ...............................50

Figure 55. Equation. Vertical displacement of a GRS abutment with a strip footing. ...................50

Figure 56. Graph. Ratio of plane strain stiffness of a strip footing on top of a wall (SGRS) to 

that of a PT (SPT) for the case of constant stiffness with depth.....................................51

Figure 57. Graph. Comparison between compacted and uncompacted GRS composites .............56

Figure 58. Design service limit for uncompacted sample DC-5......................................................57

Figure 59. Graph. Comparison of compacted and uncompacted strains between the DC-1 

and DC-5 tests ..................................................................................................................58

Figure 60. Graph. Effect of bearing bed reinforcement for TF-7 and TF-8....................................59

Figure 61. Graph. Measured lateral deformation at 3,600 psf applied stress for TF-7 (no 

bearing bed reinforcement) and TF-8 (2 courses of bearing bed reinforcement) .......60

Figure 62. Graph. Measured lateral deformation at 26,600 psf applied stress for TF-7 (no 

bearing bed reinforcement) and TF-8 (2 courses of bearing bed reinforcement) .......61

Figure 63. Graph. Comparison of open-graded and well-graded backfills for TF-1 and TF-2 ....63

Figure 64. Graph. Stress-strain curves for PTs with CMUs at Tf/Sv = 3,800 psf ...........................67

Figure 65. Graph. Stress-strain curves for PTs with no CMU facing at Tf/Sv = 3,800 psf ............68

Figure 66. Graph. Capacity of GRS with no CMU facing at various reinforcement spacing 

for different Tf/Sv ratios ...................................................................................................69

Figure 67. Graph. Capacity of GRS with CMU facing at various reinforcement spacing for 

different Tf/Sv Ratios .......................................................................................................70

Figure 68. Graph. Capacity of GRS with no CMU facing at various reinforcement strength 

for different Tf/Sv ratios ...................................................................................................70

Figure 69. Graph. Capacity of GRS with CMU facing at various reinforcement strength for 

different Tf/Sv ratios .........................................................................................................71

Figure 70. Graph. Stress-strain response for TF-2 (CMU facing) and TF-3 (No CMU facing) 

with Sv = 75⁄8 inches and Tf = 2,400 lb/ft ........................................................................72

Figure 71. Stress-strain response for TF-6 (CMU facing) and TF-7 (No CMU facing) with 

Sv = 75⁄8 inches and Tf = 4,800 lb/ft .................................................................................72

Figure 72. Graph. Stress-strain response for TF-9 (CMU facing) and TF-10 (No CMU 

facing) with Sv = 151⁄4 inches and Tf = 4,800 lb/ft .........................................................73

Figure 73. Graph. Stress-strain Response for TF-12 (CMU facing) and TF-11 (No CMU 

facing) with Sv = 3-13/16 inches and Tf = 1,400 lb/ft ....................................................74

Figure 74. Graph. Stress-strain response for TF-14 (CMU facing) and TF-13 (No CMU 

facing) with Sv = 111⁄4 inches and Tf = 3,600 lb/ft .........................................................74
vii

Figure 75. Graph. Effect of CMU facing on ultimate capacity as a function of reinforcement 

spacing ..............................................................................................................................77

Figure 76. Graph. Effect of CMU facing on ultimate capacity as a function of reinforcement 

strength .............................................................................................................................77

Figure 77. Graph. Calculated confining pressure due to CMU facing at the ultimate capacity ...78

Figure 78. Graph. Comparison of predicted capacity and measured capacity................................80

Figure 79. Graph. Cumulative distribution function plot for DC and TF PTs................................81

Figure 80. Graph. Cumulative distribution function plot for all GRS composite tests..................83

Figure 81. Graph. Normalized applied stress versus strain for all PT.............................................85

Figure 82. Graph. Normalized load-deformation behavior for the DC and TF PTs up to 

5 percent vertical strain....................................................................................................86

Figure 83. Graph. Cumulative distribution function for proposed service limit pressure .............88

Figure 84. Graph. Load-deformation behavior for the Turner Fairbank PTs at low strain 

levels.................................................................................................................................90

Figure 85. Graph. Normalized load-deformation behavior for the DC and TF PTs up to 

0.5 percent vertical strain ................................................................................................91

Figure 86. Graph. PTs strictly meeting FHWA GRS abutment design specifications...................92

Figure 87. Equation. Limit state function for FOSM approach.......................................................94

Figure 88. Graph. Reliability index for lognormal R and Q ............................................................94

Figure 89. Equation. LRFD format ...................................................................................................95

Figure 90. Equation. Resistance factor using FOSM .......................................................................95

Figure 91. Equation. Coefficient of variation for factored load ......................................................95

Figure 92. Equation. Coefficient of variation for resistance............................................................96

Figure 93. Graph. Resistance factor for footings on GRS composites for different dead to 

dead plus live load ratios and target reliability indices based on PT series .................97

Figure 94. Graph. Resistance factor for footings on GRS composites for different dead to 

dead plus live load ratios and target reliability indices based on all testing to date ....98

Figure 95. Graph. AASHTO No. 8 LSDS test results (DC tests)..................................................106

Figure 96. AASHTO No. 8 LSDS deformation test results (DC tests).........................................107

Figure 97. Graph. AASHTO No. 8 pea gravel LSDS test results (DC tests) ...............................108

Figure 98. Graph. AASHTO No. 8 pea gravel LSDS deformation test results (DC tests) ..........108

Figure 99. Graph. AASHTO No. 57 LSDS test results (DC tests)................................................109

Figure 100. Graph. AASHTO No. 57 LSDS deformation test results (DC tests) ........................110

Figure 101. Graph. AASHTO No. 9 LSDS test results (DC tests)................................................111

Figure 102. Graph. AASHTO No. 9 LSDS deformation test results (DC tests) ..........................111

Figure 103. Graph. AASHTO No. 8 LSDS test results (TFHRC tests)........................................112

Figure 104. Graph. AASHTO No. 8 LSDS deformation test results (DC tests) ..........................113

Figure 105. Graph. AASHTO A-1-a (VDOT 21A) LSDS test results (TFHRC tests)................114

Figure 106. Graph. AASHTO A-1-a (VDOT 21A) LSDS deformation test results (DC tests)..114

Figure 107. Illustration. Instrumentation layout for DC tests and TF-1........................................125

Figure 108. Illustration. Instrumentation layout for TF-2, TF-9 ...................................................125

Figure 109. Illustration. Instrumentation layout for TF-3, TF-4 ...................................................126

Figure 110. Illustration. Instrumentation layout for TF-5, TF-7 ...................................................126

Figure 111. Instrumentation layout for TF-6, TF-12......................................................................127

Figure 112. Illustration. Instrumentation layout for TF-8..............................................................127

Figure 113. Illustration. Instrumentation layout for TF-10............................................................128
viii

Figure 114. Illustration. Instrumentation layout for TF-11............................................................128

Figure 115. Illustration. Instrumentation layout for TF-13............................................................129

Figure 116. Illustration. Instrumentation layout for TF-14............................................................129
ix

LIST OF TABLES

Table 1. Summary of PT conditions..................................................................................................11

Table 2. PT reinforced backfill gradations........................................................................................12

Table 3. PT backfill gradation properties..........................................................................................13

Table 4. Maximum dry density for PT aggregates...........................................................................15

Table 5. LSDS testing results.............................................................................................................16

Table 6. Geosynthetic reinforcement properties...............................................................................17

Table 7. PT dimensions......................................................................................................................19

Table 8. PT measured results summary.............................................................................................29

Table 9. Parametric study on aggregate size.....................................................................................53

Table 10. Effect of aggregate type results.........................................................................................54

Table 11. Parametric study on compaction .......................................................................................55

Table 12. Parametric study on bearing bed reinforcement ..............................................................59

Table 13. Parametric study on gradation (Tf = 2,400 lb/ft, Sv = 75⁄8 inches)...................................62

Table 14. Parametric study on gradation (Tf = 4,800 lb/ft, Sv = 75⁄8 inches)...................................64

Table 15. Parametric study on reinforcement strength with open-graded aggregates ...................64

Table 16. Parametric study on reinforcement strength with well-graded aggregates ....................64

Table 17. Parametric study for 3,800 lb/ft2 Tf/Sv ratio (with facing) ..............................................65

Table 18. Parametric study for 3,800 lb/ft2 Tf/Sv ratio (with no facing).........................................66

Table 19. Tf/Sv ratios for each PT......................................................................................................66

Table 20. Effect of CMU facing on stiffness and capacity ..............................................................75

Table 21. Effect of CMU facing on strain.........................................................................................76

Table 22. PTs meeting GRS strength and service limit design criteria...........................................79

Table 23. Predicted and measured vertical capacity for DC and TF PTs .......................................82

Table 24. Predicted and measured vertical capacity for all GRS composite tests..........................83

Table 25. Estimation of allowable dead load to limit vertical strain to 0.5 percent using the 

GRS capacity equation......................................................................................................89

Table 26. Statistics for dead and live loads.......................................................................................96

Table 27. AASHTO No. 8 sieve analysis (DC tests) .....................................................................103

Table 28. AASHTO No. 8 pea gravel sieve analysis (DC tests) ...................................................103

Table 29. AASHTO No. 57 Sieve analysis (DC tests)...................................................................104

Table 30. AASHTO No. 9 Sieve analysis (DC tests).....................................................................104

Table 31. AASHTO No. 8 Sieve analysis (TFHRC tests) .............................................................105

Table 32. AASHTO A-1-a (VDOT 21A) sieve analysis (TFHRC tests)......................................105

Table 33. Summary of AASHTO No. 8 LSDS results (DC tests).................................................106

Table 34. Summary of AASHTO No. 8 pea gravel LSDS results (DC tests) ..............................107

Table 35. Summary of AASHTO No. 57 LSDS results (DC tests)...............................................109

Table 36. Summary of AASHTO No. 9 LSDS results (DC tests).................................................110

Table 37. Summary of AASHTO No. 8 LSDS results (TFHRC tests).........................................112

Table 38. Summary of AASHTO A-1-a (VDOT 21A) LSDS results (TFHRC tests).................113

Table 39. TF-2 Nuclear density test results ....................................................................................115

Table 40. TF-2 Nuclear density test results ....................................................................................116

Table 41. TF-3 Nuclear density test results ....................................................................................117

Table 42. TF-4 Nuclear density test results ....................................................................................117

Table 43. TF-5 Nuclear density test results ....................................................................................118
x

Table 44. TF-6 Nuclear density test results ....................................................................................118

Table 45. TF-7 Nuclear density test results ....................................................................................119

Table 46. TF-8 Nuclear density test results ....................................................................................119

Table 47. TF-9 Nuclear density test results ....................................................................................120

Table 48. TF-10 Nuclear density test results ..................................................................................120

Table 49. TF-11 Nuclear density test results..................................................................................121

Table 50. TF-12 Nuclear density test results ..................................................................................122

Table 51. TF-13 Nuclear density test results ..................................................................................123

Table 52. TF-14 Nuclear density test results ..................................................................................123

Table 53. DC-1 PT Data...................................................................................................................131

Table 54. DC-2 PT Data...................................................................................................................132

Table 55. DC-3 PT Data...................................................................................................................133

Table 56. DC-4 PT Data...................................................................................................................134

Table 57. DC-5 PT Data...................................................................................................................135

Table 58. TF-1 PT Data....................................................................................................................136

Table 59. TF-2 PT Data....................................................................................................................137

Table 60. TF-3 PT Data....................................................................................................................137

Table 61. TF-4 PT Data....................................................................................................................138

Table 62. TF-5 PT Data....................................................................................................................139

Table 63. TF-6 PT Data....................................................................................................................140

Table 64. TF-7 PT Data....................................................................................................................141

Table 65. TF-8 PT Data....................................................................................................................142

Table 66. TF-9 PT Data....................................................................................................................143

Table 67. TF-10 PT Data..................................................................................................................144

Table 68. TF-11 PT Data..................................................................................................................145

Table 69. TF-12 PT Data..................................................................................................................146

Table 70. TF-13 PT Data..................................................................................................................147

Table 71. TF-14 PT Data..................................................................................................................148
xi

LIST OF ABBREVIATIONS AND SYMBOLS

Abbreviations

AASHTO American Association of State Highway and Transportation Officials

CMU Concrete masonry unit

DC Defiance County, OH

EDC Every Day Counts initiative

FHWA Federal Highway Administration

GP-GM Poorly graded-silty gravel

GRS Geosynthetic reinforced soil

IBS Integrated bridge systems

LFD Load factor design

LRFD Load and resistance factor design

LSDS Large scale direct shear

LVDT Linear voltage displacement transducers

MARV Minimum average roll value

POT Potentiometer

PT Performance test

SRW Segmental retaining wall

TFHRC Turner-Fairbank Highway Research Center

USCS Unified Soil Classification System

Symbols

Reliability index

Slope angle

Target reliability index

β

βs

βT
xii

Unit weight of the backfill

γb

Bulk unit weight of the facing block

Maximum dry density

Load factor for dead load

Load factor for live load

Unit weight of the GRS composite

Load factor for load component i

Interface friction angle between the geosynthetic and the facing element for a 

frictionally connected GRS composite

Measured vertical strain at an applied load of 4000 psf

Measured vertical strain at failure

Maximum recorded vertical strain

Vertical strain

Vertical strain for a compacted GRS composite

Vertical strain for an uncompacted GRS composite

Bias, ratio of measured to predicted

Bias factor for dead load

Bias factor for live load

Bias factor for resistance

Poisson’s ratio of the GRS

Vertical displacement

Applied normal stress

External confining stress due to the facing

γ

γd

D

L

GRS i

δ

ε@q=4000psf ε@qult εmax v

εv,compact εv,uncompact λ

D

L

R

GRS ρ



σc
xiii

Total lateral stress within the GRS composite at a given depth and location

Shear strength of soil

Peak friction angle

Friction angle of the GRS composite

Resistance factor

Φcap

Resistance factor for capacity

Optimum moisture content

a Footing offset from the edge of the wall face (i.e., setback distance)

b Footing width on top of the GRS composite

B Base width of the GRS composite

Btotal Total width of the PT with the CMU facing

c Cohesion of the backfill

cGRS Cohesion of the GRS composite

Cc Coefficient of Curvature

Cu Coefficient of Uniformity

d Depth of the facing block unit perpendicular to the wall face

dmax Maximum aggregate size

D10 Aggregate size in which 10 percent of the sample is finer

D30 Aggregate size in which 30 percent of the sample is finer

D60 Aggregate size in which 60 percent of the sample is finer

D85 Aggregate size in which 85 percent of the sample is finer

Eo Initial stress-strain ratio

Eo,CMU Initial stress-strain ratio for tests with CMU facing

Eo, no CMU Initial stress-strain ratio for tests without any facing

σh



φ

GRS Φ

ω
xiv

EGRS Young’s modulus of the GRS composite

ER Ratio of stress to strain for the reload cycle

ḡ Mean safety margin

H Height of the GRS composite

Kar Coefficient of active earth pressure for the backfill

Kpr Coefficient of passive earth pressure for the backfill

L Length of footing/bearing area

Bearing capacity factor

Ncq Bearing capacity factor

Ns Stability factor

SGRS Plane strain stiffness of a strip footing on top of GRS

SPT Stiffness of the unconfined GRS column

Sv Reinforcement spacing

Tf Wide width tensile strength of the geosynthetic, expressed as the minimum 

average roll value (MARV)

Treq,c Required reinforcement strength in the direction perpendicular to the wall face

q Applied stress

Applied stress at 0.5 percent vertical strain

Predicted applied stress at 0.5 percent vertical strain

Applied stress at 5 percent vertical strain

qmax Maximum applied pressure during testing

qult,an,c Ultimate capacity using semi-empirical theory

qult,emp Measured failure pressure

qult,emp CMU Measured failure pressure for tests with CMU facing 

qult,emp no CMU Measured failure pressure for tests without any facing

Nq

q@ε=0.5% q@ε=0.5%, predicted q@ε=5%
xv

qult,PS Ultimate capacity of strip footing under plane strain conditions

qult,PT Ultimate capacity of the GRS column

Q Load

QD Dead load

QL Live load

Qi Load component i

R Resistance

Vdmax Coefficient of variation of the maximum aggregate size

VD Coefficient of variation of the dead load

VKp Coefficient of variation of the coefficient for passive earth pressure

VL Coefficient of variation of the live load

VM Coefficient of variation of the model

VQ Coefficient of variation of the loads

VR Coefficient of variation of the resistance 

VTf Coefficient of variation of the reinforcement strength

W Factor accounting for the effect of reinforcement spacing and aggregate size

z Standard normal variable

1

1. INTRODUCTION

The Federal Highway Administration (FHWA) has developed a standard test method to describe 

the load-deformation behavior of a frictionally connected geosynthetic reinforced soil (GRS) 

composite material which can be used to predict performance of a GRS abutment.

(1) The GRS 

performance test (PT), also called a mini-pier experiment, consists of constructing alternating 

layers of compacted granular fill and geosynthetic reinforcement with a facing element that is 

frictionally connected, then axially loading the GRS mass while measuring deformation to 

monitor performance. This large element load test provides material strength properties of a 

particular GRS composite built with different combinations of reinforcement, compacted fill, and 

facing elements. This report describes the procedure and provides axial load versus deformation 

results for a series of PTs conducted in both Defiance County, OH, as part of the FHWA’s Every 

Day Counts (EDC) GRS Validation Sessions and in McLean, VA, at the FHWA’s TurnerFairbank Highway Research Center (TFHRC) as part of a parametric study.

1.1 BACKGROUND

The FHWA first demonstrated the concept of a mini-pier experiment in 2000 with the “Vegas 

Mini-Pier” experiment, shown in figure 1.

(1,2) The pier was 8 ft tall with square inside 

dimensions of 3.5 ft, which represents a height (H) to base width (B) ratio of 2 (figure 2

through figure 4). This H/B ratio is consistent with typical triaxial testing for soils.(3)

Figure 1. Photo. Vegas mini-pier experiment.
2

Figure 2. Illustration. Plan view of Vegas mini-pier experiment.

Figure 3. Illustration. Face view of Vegas mini-pier experiment.

 

 

 

 

 

 
3

Figure 4. Illustration. Side view of Vegas mini-pier experiment.

The materials used for the GRS pier were a poorly graded-silty gravel (GP-GM, according to the 

Unified Soil Classification System, or USCS) soil with a 2,400 lb/ft (ultimate wide width tensile 

strength) geotextile spaced every 6 inches frictionally connected to segmental retaining wall 

(SRW) blocks for the facing. The top two courses of block, 1 ft from the top of the mini-pier, had

two intermediate bearing bed reinforcement layers as shown in figure 5. The resulting pier was 

loaded up to 146 psi; however, time constraints and stroke limitations for the jacks prevented 

loading to failure of the composite. Since then, several additional performance tests have been 

completed, with the largest load carrying capacity reported at 176 psi.

(4,5)
4

Figure 5. Illustration. Reinforcement schedule for Vegas mini-pier experiment.

The concept of testing GRS material has been previously applied on smaller scale models 

ranging from small triaxial sized samples to 2-ft cubed specimens in smaller capacity test 

frames.(6,7)Several large scale tests have also been conducted.

(7,8,9)

For the aggregates 

recommended by FHWA for bridge support, large scale tests are required to adequately predict 

performance of a full-scale GRS abutment.(1) The proposed FHWA PT has been shown to 

accurately predict both the strength limit and the service limit for GRS abutments.(5)

1.2 CURRENT PERFORMANCE TESTS

To investigate GRS material further, the FHWA conducted a series of 19 PTs as part of this 

research. The layout for these PTs was a slight modification of the Vegas Mini-Pier experiment. 

Since these tests were conducted with concrete masonry units (CMU) for the facing, as opposed 

to SRW blocks used in the Vegas PT, different test dimensions were needed to retain the H/B 

ratio of 2 throughout testing. For CMUs, the typical performance test is 6.4 ft tall with square 

inside dimensions of 3.2 ft (figure 6 and figure 7). The parameters that varied among tests were

reinforcement spacing (from 4 to 16 inches), geotextile strength (from 1,400 to 4,800 lb/ft, soil 

type (open-graded and well-graded), and frictionally connected facing element (concrete 

masonry unit facing and no facing).

 

 

 
5

Figure 6. Illustration. Plan view of Defiance County experiment.

Figure 7. Illustration. Elevation view of Defiance County (DC) test.

1.3 INTERNAL STABILITY DESIGN

The results of the PT are primarily used in the design of GRS abutments.

(1) The resulting stressstrain curve can be used to describe the strength limit state for capacity and the service limit state 

for deformation (vertical and lateral) due to an applied load. It is the only method currently 

available to describe both the capacity and deformation behavior of GRS for load bearing 

applications; American Association of State Highway and Transportation Officials (AASHTO)

(2012) does not provide any guidance for these GRS limit states.(10)

 

 

 

 

 

 
6

In addition to the empirically based design using PT results, a semi-empirical method is also 

presented by FHWA for internal stability design of GRS abutments and integrated bridge 

systems, based on basic soil mechanics principles for the relationship between vertical and lateral 

stress but is fit to the available data at the time.

(1) The ultimate capacity is determined according 

to the equation presented in figure 8.

(11) Note that in design, the effect of confining stress (due to 

the facing element) and cohesion is ignored, which for a CMU facing and the select granular fill 

specified for GRS abutments is an appropriate assumption.(5)

Figure 8. Equation. Ultimate vertical capacity for a GRS composite.

Where qult,an,c is the ultimate capacity using semi-empirical theory, is the external confining 

stress due to the facing, W is a factor accounting for the effect of reinforcement spacing and 

aggregate size (figure 9), Tf is the minimum average roll value (MARV) for wide width tensile 

strength of the geosynthetic, Kpr is the coefficient of passive earth pressure for the backfill, and c

is the cohesion of the backfill.

Figure 9. Equation. W factor.

Where Sv is the reinforcement spacing and dmax is the maximum aggregate size. The constant of 

0.7 in the W-factor is the ratio of average force in the reinforcement to the maximum force in the 

reinforcement and the 6dmax term is an approximation of the minimum distance between 

reinforcement layers where the grain size of the soil will not significantly impact the composite 

behavior of the GRS.(12)

A variant of figure 8 is also used to determine the required reinforcement strength at a given 

applied stress (figure 10). The method provides engineers with the minimum strength of 

reinforcement needed to prevent failure of the GRS composite at that given applied stress.

Note that figure 10 will not represent the actual load on the reinforcement at working stress

once increased reinforcement strengths are used in the field since this will change the conditions.

The actual load on the reinforcement is difficult to estimate; instead, the service limit for the 

reinforcement is based on limiting reinforcement strain to 2 percent.

(1)

Figure 10. Equation. Required reinforcement strength.

Where Treq,c is the required reinforcement strength in the direction perpendicular to the wall face, 

is the total lateral stress within the GRS composite at a given depth and location, is the 

external confining stress due to the facing, c is the cohesion of the reinforced backfill, Kar is the 

coefficient of active earth pressure for the backfill, W is a factor accounting for the effect of 

qult,an,c = σc + W 

Tf

Sv

 Kpr + 2c Kpr

σc

W = 0. 7

 

Sv

6dmax

 

Treq,c = 

σh − σC − 2c Kar

W

 Sv

σh σc
7

reinforcement spacing and aggregate size (figure 9), and Sv is the reinforcement spacing. As with 

the ultimate capacity equation, the effect of confining stress (due to the facing element) and 

cohesion is ignored when determining the required reinforcement strength in the design of GRS 

abutments, leading to added conservatism in the design method.

(5)

The soil-geosynthetic composite capacity equation (figure 8) and the required reinforcement 

strength equation (figure 10) have been previously validated against the results of 16 different 

types of tests, including previous performance tests (figure 12 and figure 13, respectively). The 

research study presented in this report will add to this database to further quantify the predictive 

capability of these equations.

Note that the predictions for capacity and required reinforcement strength used the estimated

confining stress calculated according to figure 11, and the measured cohesion from soil testing,

both of which would be ignored in design. In addition, the minimum average roll value (MARV)

for the reinforcement, which is a property value calculated as the average strength for a roll less 

two standard deviations, was used to calibrate the predictive equations since the actual strength 

of the reinforcement for each test was not measured or precisely known, and the MARV is an 

industry value common for all geosynthetics specified in practice.

Figure 11. Equation. Confining stress.

(11)

Where is the bulk unit weight of the facing block, d is the depth of the facing block unit 

perpendicular to the wall face, and is the interface friction angle between the geosynthetic and 

the facing element for a frictionally connected GRS composite.

Note that the friction angle and cohesion terms used to calculate the earth pressure coefficients 

and lateral stress within the ultimate capacity (figure 8) and required reinforcement strength 

(figure 10) equations were measured using a large scale direct shear (LSDS) test on the backfill 

alone according to ASTM D3080.

(13) The peak strength of the reinforced backfill material was 

selected as it is a commonly reported value and easiest to ascertain from typical testing. More 

details are provided in section 2.1.3.

σc = γbdtanδ

γb

δ
8

Figure 12. Graph. Predictive capability of the soil-geosynthetic 

composite capacity equation.(5)

Figure 13. Graph. Predictive capability of the required reinforcement strength equation.(5)
9

1.4 OBJECTIVES

The results of the PT parametric study will be used for several purposes. The primary objectives 

of this research report are to: (1) build a database of GRS material properties that can be used by 

designers or for Load and Resistance Factor Design (LRFD) calibration; (2) evaluate the 

relationship between reinforcement strength and spacing; (3) quantify the contribution of the 

frictionally connected facing elements at the service limit and strength limit states; (4) assess the 

new internal stability design method proposed by Adams et al. 2011a for GRS; and (5) perform a 

reliability analysis of the proposed soil-geosynthetic capacity equation for LRFD.(1,11) When 

combining additional measurement techniques (e.g., contact pressure cells, earth pressure cells, 

etc.), other uses for the PT are possible, such as to study thrust against the face as a function of 

spacing; however, this is outside the scope of this report.

11

2. TESTING CONDITIONS

A series of 19 performance tests have been conducted (table 1); 5 at the Defiance County, OH,

highway maintenance facility and 14 at the TFHRC.

Table 1. Summary of PT conditions.

Test

No.

Backfill Reinforcement

Facing

Type

(°)

c

(psf)

dmax

(inch)

Tf

^

(lb/ft)

Sv

(inch)

Tf/Sv

(lb/ft2)

DC-1 8 54 0 1⁄2 4,800 75⁄8** 7,600 CMU

DC-2 8P* 46 0 3⁄4 4,800 75⁄8** 7,600 CMU

DC-3 57 52 0 1 4,800 75⁄8** 7,600 CMU

DC-4 9 49 0 3⁄8 4,800 75⁄8** 7,600 CMU

DC-5 8*** 54 0 1⁄2 4,800 75⁄8** 7,600 CMU

TF-1

++ 8 55 0 1⁄2 2,400 75⁄8 3,800 CMU

TF-2 21A 53 115 1 2,400 75⁄8 3,800 CMU

TF-3 21A 53 115 1 2,400 75⁄8 3,800 no CMU

TF-4

+

21A 53 115 1 4,800 75⁄8 7,600 no CMU

TF-5

++ 21A 53 115 1 4,800 75⁄8 7,600 no CMU

TF-6

++ 21A 53 115 1 4,800 75⁄8 7,600 CMU

TF-7 21A 53 115 1 4,800 75⁄8 7,600 no CMU

TF-8 21A 53 115 1 4,800 75⁄8** 7,600 no CMU

TF-9 21A 53 115 1 4,800 151⁄4 3,800 CMU

TF-10 21A 53 115 1 4,800 151⁄4 3,800 no CMU

TF-11 21A 53 115 1 1,400 3

13/16 4,400 no CMU

TF-12 21A 53 115 1 1,400 3

13/16 4,400 CMU

TF-13 21A 53 115 1 3,600 111⁄4 3,800 no CMU

TF-14 21A 53 115 1 3,600 111⁄4 3,800 CMU

φ = the peak friction angle, c = the cohesion at peak strength, dmax = the maximum aggregate 

size, Tf = the ultimate reinforcement strength, expressed as the minimum average roll value 

(MARV) from ASTM D4595 testing,(1)and Sv = the reinforcement spacing.

^ MARV value.

*Rounded pea-gravel angularity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

***Uncompacted sample, +technical difficulties required termination during testing.

++Technical difficulties resulted in unloading/reloading of the composite.

2.1 BACKFILL CONDITIONS

In total, six unique backfill types were used in the PTs: (1) an AASHTO No. 8 crushed, 

manufactured limestone aggregate obtained from Defiance County OH, (2) an AASHTO No. 8 

rounded quartz pea gravel (PG) obtained from Defiance County, OH, (3) an AASHTO No. 57 

crushed, manufactured limestone aggregate obtained from Defiance County, OH, (4) an 

AASHTO No. 9 crushed, manufactured limestone aggregate obtained from Defiance County, 

φ
12

OH, (5) an AASHTO No. 8 crushed, manufactured diabase aggregate obtained from Loudon 

County, VA, and (6) an AASHTO A-1-a aggregate obtained from Loudon County, VA (also 

referred to locally as a Virginia DOT, or VDOT, 21A material).

2.1.1 Sieve Analysis

All backfills used for testing, except for the AASHTO No. 9 aggregates, meet the FHWA 

specifications for use in bridge abutments.

(1) The gradations of each aggregate are shown in 

table 2 and figure 14.

Table 2. PT reinforced backfill gradations.

Sieve 

No.

Percent Passing

No. 8

(OH)

No. 8 PG

(OH)

No. 57

(OH)

No. 9

(OH)

No. 8

(VA)

A-1-a

(VDOT 21A)

1.5 100.00 100.00 100.00

1 100.00 100.00 100.00

0.75 100.00 100.00 87.91

0.50 100.00 99.57 35.69 99.69 82.41

0.375 96.99 95.58 14.13 100.00 69.86 71.36

4 26.50 14.60 3.67 94.22 7.78 48.52

8 4.63 6.98 2.41 27.75 1.66 35.24

10 1.39 32.81

16 1.85 4.34 1.47 8.90 1.11 25.40

40 0.93 16.66

50 0.91 2.64 3.66 0.88

100 0.76 3.22

200 0.65 0.71 2.82 6.47

Blank cell = no value was measured for that particular sieve number.

Most of the backfills tested are open or poorly graded materials (e.g., AASHTO Nos. 57, 8, 

and 9); however, the A-1-a material is a well-graded material. Table 3 shows the classification 

(based on the USCS) along with the maximum aggregate size (dmax), other relevant grain sizes 

for various percent passing values, and the coefficient of curvature (Cc) and coefficient of 

uniformity (Cu) for each material tested.
13

Table 3. PT backfill gradation properties.

Aggregate

Type

USCS

Classification

dmax

(inch)

D85

(inch)

D60

(inch)

D30

(inch)

D10

(inch) Cc Cu

8 (OH) GP 0.50 0.34 0.28 0.20 0.12 1.19 2.36

8 PG (OH) GP 0.75 0.35 0.29 0.22 0.13 1.30 2.23

57 (OH) GP 1.00 0.74 0.62 0.47 0.30 1.18 2.05

9 (OH) SP 0.38 0.17 0.14 0.47 0.05 1.35 2.78

8 (VA) GP 1.00 0.44 0.35 0.25 0.19 0.96 1.78

A-1-a

(VDOT 21A)

GW-GM 1.00 0.57 0.28 0.07 0.01 2.67 46.67

dmax = the maximum aggregate size.

D85 = the aggregate size in which 85 percent of the sample is finer.

D60 = the aggregate size in which 60 percent of the sample is finer.

D30 = the aggregate size in which 30 percent of the sample is finer.

D10 = the aggregate size in which 10 percent of the sample is finer.

Cc = the coefficient of curvature.

Cu = the coefficient of uniformity.
14

Figure 14. Graph. Reinforced backfill gradations.
15

2.1.2 Density

To determine the in-place compaction requirements of the backfill material, a standard Proctor 

test was conducted according to Method D of AASHTO T99 for the AASHTO A-1-a (VDOT 

21A) material.(15)In addition, vibratory tests were conducted according to ASTM D4252 for the 

open-graded materials to determine the maximum dry density.(16) The results are shown in 

table 4.

Table 4. Maximum dry density for PT aggregates.

Aggregate Type

Max Dry Density, 

(pcf)

Optimum Moisture Content, 

(percent)

8 (OH) 101.27 N/A

8 PG (OH) 115.75 N/A

57 (OH) 108.69 N/A

9 (OH) 110.66 N/A

8 (VA) 112.82 N/A

A-1-a (VDOT 21A) 148.90 7.7

N/A = Not applicable, there is no optimum moisture content for open-graded aggregates 

since they are free draining.

2.1.3 Friction Angle

The strength properties of each aggregate were determined using a large scale direct shear 

(LSDS) device according to ASTM D3080.

(13) The LSDS device at TFHRC is 12 x 12 x 8 inches

in dimension and is capable of testing aggregates up to 1.2 inches. For this series of experiments, 

the unscalped aggregates were tested at four applied normal stresses, 5, 10, 20, and 30 psi, at a 

shear rate of 0.015 inches/min and a gap size equal to the D85 of the material (i.e., the aggregate 

size where 85 percent of the sample is smaller; see table 3).

The open-graded materials were tested in a dry, uncompacted state prior to the consolidation 

phase in the LSDS device. The well-graded material (VDOT 21A) was tested at 100 percent of 

the maximum dry density (i.e., the level of compaction achieved during each PT with the 

backfill), and at the optimum moisture content (see table 4). Since the shear strength failure 

envelope for these aggregates is non-linear, the reported cohesion for the well-graded material 

(VDOT21A) was determined through a series of LSDS tests that were performed with the 

compacted state fully saturated; note that the resulting peak friction angle in this case was similar 

to the non-saturated condition.

The results of the LSDS testing are shown in table 5 and in figure 15. Note that the reported 

friction angle is based on the measured peak strength during testing and assumes a linear MohrCoulomb envelope for the range of confining stresses tested. Note that peak strength for these 

backfill materials is mobilized at typically 0.5- to 1-inch lateral displacement in the LSDS 

device, which corresponds to about 4- to 8-percent lateral strain for the 12-inch shear box. In the 

context of the PT though, the GRS composite is tested to failure, sometimes well beyond 

8-percent lateral strain. From a theoretical perspective, it may be more appropriate to model the 

γd ω
16

failure of a GRS composite by using the friction angle at the fully softened state of the backfill 

material during the LSDS test; however, to conform to the current standard-of-practice, the peak 

strength of the reinforced backfill material was selected to calibrate the design as it is the 

commonly reported value and easiest to ascertain from typical testing. Figure 15 provides the 

raw data for LSDS testing.

Table 5. LSDS testing results.

Aggregate Type Friction Angle (°) Cohesion (psf)

8 (OH) 54 0

8 PG (OH) 46 0

57 (OH) 52 0

9 (OH) 53 0

8 (VA) 55 0

A-1-a (VDOT 21A) 54 115

Figure 15. Graph. LSDS testing results.

0

10

20

30

40

50

60

0 5 10 15 20 25 30 35 40

Shear Stress (psi)

Normal Stress (psi)

No. 8 (OH)

No. 8 PG (OH)

No. 57 (OH)

No. 9 (OH)

No. 8 (VA)

VDOT 21A

VDOT 21A (Saturated)
17

2.2 REINFORCEMENT CONDITIONS

In all of the tests, a biaxial, woven polypropylene geotextile was used as the reinforcement 

element; however, different strengths and stiffness of material were used among the PTs. The 

manufacturer supplied MARV data is shown in table 6. A more recent property of the 

geosynthetic used in design is the wide width tensile strength at 2-percent strain which provides 

an indication of GRS performance at the service limit state.

(1) Currently, many of the 

manufacturers do not make this value publicly available, although they can supply it on request.

Note that in the field, the actual strength of the reinforcement will be higher than the reported 

MARV.

Table 6. Geosynthetic reinforcement properties.

PTs:

TF-11, 

TF-12

TF-1, 

TF-2, 

TF-3

TF-13, 

TF-14

DC-1, DC-2, 

DC-3, DC-4, 

DC-5, TF-4, 

TF-5, TF-6, 

TF-7, TF-8, 

TF-9, TF-10

Reference: (17) (18) (19) (18)

Property Test 

Method Minimum Average Roll Value (MARV)

1

Tensile 

Strength 

(Grab)

ASTM 

D4632(20) 200 x 200 lb 315 x 300 lb 450 x 350 lb 600 x 500 lb

Wide 

Width 

Tensile

ASTM 

D4595

(14)

1,400 x 

1,400 lb/ft

2,400 x 

2,400 lb/ft

3,600 x 3,600 

lb/ft

4,800 x 

4,800 lb/ft

Wide 

Width 

Elongation

ASTM 

D4595

(14) 9 x 7 percent

10 x 8

percent 15 x 10 percent

10 x 8

percent

Wide 

Width 

Tensile 

Strength at 

5 Percent

Strain

ASTM 

D4595

(14)

Not 

specified

884 x 1,564 

lb/ft

1,392 x 1,740 

lb/ft

660 x 1,500 

lb/ft

1Values for the machine (warp) by cross machine (fill) directions, respectively.

2.3 FACING CONDITIONS

A concrete masonry unit (CMU) facing was used on 11 of the 17 tests; the remaining 6 tests 

were conducted without any facing element. The CMU is a dry-cast, split-faced product with

dimensions of 75⁄8 x 75⁄8 x 155⁄8 inches and an approximate weight of 42 lb. The CMUs are 

frictionally connected to the geotextile reinforcement. The reinforcement overlaps at least 

85 percent of the block depth, termed the coverage ratio, as specified by Adams et al. 2011a.

(1)

19

3. TEST SETUP

3.1 LAYOUT

The dimensions for each PT are given in table 7. For each test, the base to height ratio was kept

constant at about 2 to mimic triaxial conditions. The height is equivalent to 10 courses of CMU 

block, while the outside dimensions of the PT with facing are 3.5 blocks wide. For each test, the 

total width of the PT with the CMU facing (Btotal) is 541⁄2 inches; the width of the GRS composite 

itself (B) is 391⁄4 inches. The footing width on top of the GRS composite (b) is 36 inches.

Table 7. PT dimensions.

Test No No. Reinf. Layers

Sv

(inch)

H 

(inch) H/B

DC-1 9 75⁄8** 761⁄4 1.9

DC-2 9 75⁄8** 761⁄4 1.9

DC-3 9 75⁄8** 761⁄4 1.9

DC-4 9 75⁄8** 761⁄4 1.9

DC-5 9 75⁄8** 761⁄4 1.9

TF-1 9 75⁄8 761⁄4 1.9

TF-2 9 75⁄8 761⁄4 1.9

TF-3 9 75⁄8 761⁄4 1.9

TF-4 9 75⁄8 761⁄4 1.9

TF-5 9 75⁄8 761⁄4 1.9

TF-6 9 75⁄8 761⁄4 1.9

TF-7 9 75⁄8 761⁄4 1.9

TF-8 9 75⁄8** 761⁄4 1.9

TF-9 4 151⁄4 761⁄4 1.9

TF-10 4 151⁄4 761⁄4 1.9

TF-11 19 3

13/16 761⁄4 1.9

TF-12 19 3

13/16 761⁄4 1.9

TF-13 6 111⁄4 783⁄4 2.0

TF-14 6 111⁄4 783⁄4 2.0

Sv = the reinforcement spacing, H = the height of the PT, 

B = the width of the GRS composite (without the facing 

element).

**Two courses of bearing bed reinforcement placed at the top 

of the PT.
20

Figure 16. Photo. DC-1 GRS PT (before testing).

3.2 LOAD AND REACTION FRAME

The reaction frames were slightly different among the tests. The Defiance County PTs (DC-1 

through DC-5) were built on a reinforced concrete base pad similar to the Vegas mini-pier setup

(figure 3 and figure 4). The base pad was elevated on CMU blocks to make room for the bottom 

set of bolted channel beams, while the top set of bolted channels was supported on the top 

concrete pad which was centered on the GRS composite. The top load pad is supported on the 

GRS composite, inside the perimeter of the CMU facing; there is an inset of 15⁄8 inches around 

the load pad and the back of the CMU facing (figure 17). The upper and lower channel beams 

were coupled together with high strength post-tensioning bar. Four hollow core hydraulic jacks 

were bolted to the top channel beams (figure 18). All jacks were connected to a manifold and 

controlled with an electric hydraulic pump controlled by a push button solenoid valve. The 

stroke and capacity of the hollow core jacks were 6 inches and 120 kips, respectively.
21

Figure 17. Illustration. Concrete footing on GRS composite, inset from facing.

Figure 18. Photo. Hollow core hydraulic jacks for PT assembly.

A similar reaction frame setup was used for the initial TFHRC PTs (figure 19), except the 

concrete strong floor was used for the reaction instead of sandwiching the GRS composite 

between twin sets of bolted channels (figure 4). The top sets of bolted channels were placed on 

an elastomeric pad and positioned on the top concrete pad to evenly distribute the load delivered 

from the hydraulic jacks; a 1-inch steel plate was also bolted to the top concrete pad to increase

the rigidity of the concrete.
22

Figure 19. Photo. TF-1 PT setup with reaction frame.

The reaction frame was changed after eccentric loading was observed with the previous setup on 

the TF-4 test. Subsequent TF tests used a rigid two post frame that provided additional benefits 

in terms of easier setup and the ability to gang two high-capacity, 600-kips jacks to ensure 

sufficient stroke to fail any GRS composite without having to reset the loading beam position

(figure 20, figure 21, and figure 22). A spherical bearing was used in the two-post frame 

assembly to maintain a vertical load on the top bearing pad (figure 23). Figure 23 also illustrates 

that the top concrete load pad was pushed approximately 18 inches into the GRS composite 

before failure was induced.
23

Figure 20. Photo. TF-6 PT setup with reaction frame.

Figure 21. Photo. TF-10 PT setup with reaction frame.
24

Figure 22. Photo. TF-9 at failure with reaction frame. 
25

Figure 23. Photo. Spherical bearing to apply load to the footing on the GRS composite.

3.3 CONSTRUCTION

The method of construction for the mini piers was basically the same for all of the PTs; the GRS 

composite was built from the bottom up, one course of block at a time. The primary difference in 

construction between the Defiance County tests and the TFHRC tests was the level of quality 

control. The five Defiance County tests (DC-1 through DC-5) were each constructed with the 

same supervisor, but different crews, as part of the FHWA’s EDC-GRS Validation Sessions in 

2010. The quality control for these tests was therefore limited, which may have impacted the 

testing results, particularly the stiffness. The 14 PTs conducted at TFHRC were constructed in 

the Structures Laboratory and were research quality experiments, with consistent construction 

procedures and compaction control.

At the start of construction for each PT, the first course of block was placed level and centered 

within the position of the reaction assembly. Aggregate was then infilled using either a front-end 

loader (for the DC test series) or a concrete dump hopper (for the TF test series). For the TF 

tests, the squareness of the facing and the verticality of the composite were checked before 

placement of fill in each course of block by measuring the inside diagonal lengths and checking 

with a plumb bob on every other lift.

Compaction for the DC tests was performed using a standard 18-inch-wide, gas powered

vibratory plate compactor at each lift of aggregate, a nominal 8-inch thickness (height of the 

CMU block). For the TF tests, it was necessary to compact each 8-inch lift of fill in half layer 

increments, or at a lift thickness of about 4 inches because a lightweight electric vibratory plate 

compactor was used for construction; with this equipment, it was impossible to achieve the target 
26

density at the nominal 8-inch lifts. The electric compactor had a 10-square-inch plate, with a 

weight of 48 lb and an adjustable impact force of 700 lb.

For the well-graded AASHTO A-1-a (VDOT 21A) aggregate, compaction control consisted of 

testing each nominal 4-inch lift (half the height of a CMU) with a nuclear density gauge to 

ensure the target density and moisture content (± 2 percent) were met throughout the height of 

the GRS composite. In the field, density tests are typically performed after the compaction of fill 

material to the height of CMU block; however, during construction of the TF PTs using wellgraded backfill, density testing was performed on nominal 4-inch lifts to test the compaction 

effort of the lightweight (48 lb) tamper used. For each lift, approximately 100 percent of the 

standard Proctor value was achieved; the results of the nuclear density gauge for each test are

shown in appendix B.

For the open-graded aggregates, compaction was performed until no further compression of the 

aggregate was observed, which consisted of running the plate compactor across each lift with a 

minimum of four passes. While it is not common practice to test open-graded aggregates with a 

nuclear density gauge, backscatter tests were performed on each nominal 4-inch lift for PT TF-1 

to collect the surficial density value (appendix B). The percent compaction is based on the 

maximum index density determined from ASTM D4253 (table 4).

(16)

Rodding with a shovel end was also used for each PT to compact the aggregate at the corners and 

edges of the facing. Once final compaction was achieved to the leveled height of the facing block 

and before placement of the next course of CMU blocks, any remaining aggregate was brushed 

off of the facing block so that a smooth surface existed to prevent point loading and ensure even 

placement of the next layer of block. Depending on the reinforcement schedule, a layer of 

geotextile reinforcement was then placed over the aggregate with a facing element coverage ratio 

of at least 85 percent the width of facing element. While the geotextiles used in each test were 

biaxial, the stiffness in the machine (warp) and cross-machine (fill) directions for the 

reinforcement was different (table 6). For this reason, the reinforcement was placed in an 

alternating pattern with each subsequent layer to prevent preferential failure of the PT in the 

weaker reinforcement direction.

To facilitate construction, two sets of ratchet straps were placed around the facing blocks to 

secure and maintain block alignment during compaction. During each layer of GRS construction, 

the lower strap was removed and then used to band the new, upper course of block; the process 

was repeated until the pier was completed. This process was employed because of the geometry 

and the reduced footprint of the PT where lateral block movement during compaction is 

increased compared to a typical in-service bridge abutment (plane strain condition). The ratchet 

strap on the top course of block was removed after specimen construction was fully completed,

prior to loading.

Once the fabric was placed, the next layer of CMU blocks were positioned and the aggregate 

infilled. This process was repeated for 10 courses of CMU blocks. The concrete footing was then 

placed on top of the GRS composite and the load frame assembled for testing. Construction of 

the open-graded aggregate tests took about 4 hours, and 20 hours for the DC tests and TF tests, 

respectively. The principal difference in time was attributable to the need for compaction in 
27

4-inch lifts, the requirement for nuclear density testing, and the level of quality control. 

Construction of the well-graded aggregate tests at TF took approximately 30 hours.

3.4 INSTRUMENTATION

Each PT was instrumented to measure the response to the static, vertical applied pressure on the 

top of the GRS composite. Vertical and lateral deformation was constantly monitored with 

displacement transducers. Four transducers measured the vertical displacement at the middle of 

each loading block edge while five transducers measured the horizontal displacement along one 

face of the specimen (figure 24). For the five PTs conducted in Defiance County, OH, (DC-1 

through DC-5) and the first PT conducted at TFHRC (TF-1), linear voltage displacement 

transducers (LVDT) were used.

String potentiometers (POT) replaced the LVDTs for the remaining tests at TFHRC with CMU 

blocks (TF-2, TF-6, TF-9, TF-12, TF-14) to measure both vertical and lateral deformation. For 

the PTs without any facing element (TF-3, TF-4, TF-5, TF-7, TF-8, TF-10, TF-11, TF-13), 

LVDTs were still used to measure lateral deformation in addition to the string POTs used to 

measure vertical deformation. The instrumentation layouts for deformation measurements in 

each test, similar to that shown in figure 24, are presented in appendix C. With the TFHRC PT 

series, additional instrumentation was included to monitor vertical pressure near the base of the 

composite, strain on different geotextile layers, and lateral earth pressure behind the face just 

above the center height of the composite (figure 25). This report focuses on the vertical 

deformation data, which can be immediately used in the design of GRS composites for load 

bearing applications.

(1)

Figure 24. Illustration. Instrumentation layout for DC tests and TF-1.
28

Figure 25. Illustration. General additional instrumentation layout TF PT series.

3.5 LOAD SCHEDULE AND COLLECTION OF DATA

For each TFHRC test, the loading schedule was prepared in the same manner. The basic protocol 

included estimating the ultimate vertical capacity determined from the semi-empirical equation 

(figure 8). The estimated capacity was then divided into a minimum of ten load increments to

capture the deformation characteristics of each test. In addition, between two and four load 

increments below the 4,000 psf service limit were included in each test to better define the 

behavior of the composite at low applied loads.

Each load increment was applied manually using an electric hydraulic pump with a push button 

controlled solenoid valve. 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 operation of the system throughout the course of the load test. In each 

experiment, load was increased, from increment to increment, only when there was no significant 

change (< 0.003 inch) in settlement between any two time intervals; 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 settlement at 1-, 3-, 5-, 7-, 15-, 20-, 25-, and 30-minute

intervals from the start of each load increment.

The settlement of the top footing, lateral displacement of the face, and geotextile strain were 

recorded every minute within each load increment. Any vertical earth pressure measurements

and lateral stress information were collected at the end of each load increment. Photographs were 

also taken at the conclusion of each load increment to visually document the test. Typically, each 

test took about 6 hours to complete.
29

4. RESULTS DATABASE

The results for each test are presented in table 8. The corresponding vertical stress-strain curves 

are shown in figure 26 and figure 27 for the Defiance County tests (labeled DC) and the TFHRC 

tests (labeled TF), respectively. The tabular data for each stress-strain curve shown in figure 26

and figure 27 is given in appendix D. Note that the applied stress and strain are calculated as 

averages of the measured values (over the period of loading for stress and over the four

LVDT/POTs located on the footing for strain at the end of each load increment).

Table 8. PT measured results summary.

Maximum Tested

Strength 

Limit Design Limit

Service 

Limit

Test

No.

qmax

(psf) (%)

qult,emp

(psf) (psf)

Vallow,emp = 

/3.5 

(psf)

(psf)

DC-1 23,310 7.95 23,310 19,983 5,709 3,065

DC-2 22,709 7.07 22,709 19,399 5,543 2,171

DC-3 18,447 5.82 N/A 16,182 4,623 1,324

DC-4 26,730 7.64 N/A 17,350 4,957 2,212

DC-5 21,539 10.42 21,539 11,761 3,360 316

TF-1 20,487 10.9 20,487 13,409 3,831 2,075

TF-2 25,260 11.46 25,260 18,711 5,346 4,759

TF-3 17,491 13.8 17,491 12,120 3,463 3,417

TF-4 14,240 4.4 N/A N/A1 N/A13,705

TF-5 25,920 17.9 25,920 15,581 4,452 1,815

TF-6 43,828 15.7 43,828 22,007 6,288 3,704

TF-7 26,546 12.5 26,546 13,684 3,910 2,224

TF-8 29,134 17.8 29,134 13,797 3,942 1,753

TF-9 22,310 15.6 22,310 13,527 3,865 2,955

TF-10 10,330 14.27 10,330 7,374 2,107 1,586

TF-11 23,249 12.79 23,249 13,316 3,805 2,839

TF-12 29,030 13.37 29,030 18,573 5,307 4,028

TF-13 12,960 12.32 12,960 8,641 2,469 2,398

TF-14 23,562 12.69 23,562 16,748 4,785 2,037

qmax = the maximum applied pressure during testing, = the maximum recorded vertical 

strain, qult,emp is the measured failure pressure, = the applied stress at 5 percent 

vertical strain, Vallow,emp is the total allowable pressure on the GRS,

(1)

= the applied 

stress at 0.5 percent vertical strain.

N/A = Not applicable, did not fail composite. N/A1= Not applicable, composite did not reach 

5-percent vertical strain before it was terminated.

εmax q@ε=5% q@ε=5% q@ε=0.5% εmax q@ε=5% q@ε=0.5%
30

Figure 26. Graph. Load-deformation behavior for the Defiance County PTs.
31

Figure 27. Graph. Load-deformation behavior for the Turner Fairbank PTs.
32

Note that two Defiance County (DC-3 and DC-4) and one TFHRC test (TF-4) were prematurely 

terminated before failure of the GRS composite. For the DC tests, the 6-inch stroke of the 

hydraulic jacks limited the amount of deformation that could be applied to the samples without 

resetting the experiment, which did not occur. For the TF test, the applied load on TF-4 was not 

uniform, leading to tilting of the footing (figure 28). Analysis of the data shows this to be the 

case.

 

Figure 28. Photo. Tilting of the footing during TF-4 testing.

The TF-4 data plotted on the family of load test curves shown in figure 27 is the average 

measured settlement from the four string POTs located on each side of the top of the footing. 

Figure 29 shows the settlement data for each POT, along with the average, for the TF-4 test. As 

indicated visually (figure 28), the measurements (figure 29) confirm that the footing titled 

towards sides 3 and 4 of the footing. Figure 28 also illustrates the composite behavior of GRS 

(i.e., the soil and geosynthetic tend to move together as one material).
33

Figure 29. Graph. TF-4 results.

4.1 UNLOAD/RELOAD BEHAVIOR

TF-1, TF-5 and TF-6 experienced technical difficulties during testing that resulted in an 

unloading/reloading of the composite (figure 30, figure 31 and figure 32, respectively). The load 

test curves shown in figure 27 were corrected to omit the unload/reload cycles.

Due to a power outage at TFHRC during testing for TF-1 (figure 30), the GRS composite was 

unloaded for a period of over 40 hours (1.7 days); the maximum applied stress prior to unloading 

of the composite was 12,761 psf. After reestablishing power, the sample was then reloaded with 

the approximate same loading schedule as the initial load until failure of the composite was 

reached at 20,487 psf. The initial ratio of stress to strain (Eo) is equal to approximately 320 ksf,

while the reload response was considerably stiffer by a factor of over 6, with a reload ratio of 

stress to strain (ER) around 2,070 ksf.

For TF-5 (figure 31), the GRS composite was initially loaded and reloaded twice at low applied 

loads (less than 4,700 psf). The string potentiometers were reset in each case. An additional 

unloaded cycle then occurred for a period of 12 minutes 50 seconds in order to reset the single 

hollow-core jack because of the stroke limitation (hence, the change to the use of two jacks in 

series starting with TF-6). After resetting the jack, the sample was then reloaded at the 

approximate same loading schedule as the initial load. The last reload cycle that occurred during
34

TF-5 was at an applied stress of about 25,000 psf (14th load increment). The failure load was 

measured at 25,920 psf meaning the unload/reload cycle occurred near failure when the GRS 

composite had already experienced considerable spalling and necking because of the lack of the 

facing element. The initial ratio of stress to strain for virgin loading (Eo) of TF-5 during the first 

applied loading sequence is about 560 ksf. The first and second reload cycles were considerably 

stiffer, both with a reload ratio of stress to strain equal to approximately 2,900 ksf, supporting the 

shakedown concept. The last reload cycle; however, is not as stiff as the first two. It is 

hypothesized by the authors that the strain hardening effect observed in TF-1 was not as 

pronounced during TF-5 at higher applied loads because the GRS composite had necked 

considerably prior to the last reload cycle. TF-7 was therefore conducted to determine if there 

was an effect of the necking of the sample on the capacity of the composite.

The two unload/reload cycles that occurred during TF-6 were at applied stresses of 

approximately 4,500 psf and 8,000 psf (5th and 10th load increment, respectively). Figure 32

indicates that preloading of the GRS composite resulted in strain hardening to produce a stiffer 

response during the reload cycle. The reload response is similar during both cycles, with a ratio 

of stress to strain (ER) equal to approximately 3,000 ksf, compared to the ratio for the initial, 

virgin loading (Eo) equal to approximately 750 ksf, an increase by a factor of four.

Typical values of the soil modulus for shallow foundation applications is between about 200 to 

400 ksf for natural foundation soils.(21) Based on the results of this series of PTs, GRS 

composites are stiffer, with modulus values between about 320 and 750 ksf. Preloading results in 

an even stiffer composite. The strain hardening effect due to preloading was observed during the 

FHWA GRS Pier load test in 1995 and in-service, with the preloaded GRS abutments for the 

Blackhawk bridge built in Central City, CO, in 1998.

(22,23) To study the possible shakedown 

phenomenon and resulting response under repeated unload/reload cycles, additional performance 

testing under cyclic load is needed.
35

Figure 30. Graph. TF-1 results.
36

Figure 31. Graph. TF-5 results.
37

Figure 32. Graph. TF-6 results.
38

4.2 REPEATABILITY

The repeatability of the test method was indirectly measured through four PTs (TF-4, TF-5, 

TF-7, and TF-8, see TABLE 1) on nearly identical GRS composites (figure 33). Note that TF-4 

was prematurely terminated due to uneven loading and settlement and TF-8 had two courses of 

bearing bed reinforcement at the top. For TF-5, the stress-strain curve for both the initial loading 

sequence is plotted to compare to the other tests at lower applied loads, with the stress-strain 

curve starting on the second reload cycle also plotted to compare to the other tests at higher 

applied loads. All of the tests were performed without CMU facing. The results indicate good 

agreement and suggest that the performance test produces repeatable results.

4.3 COMPOSITE BEHAVIOR

The new internal stability design method proposed by Adams et al. 2011a for GRS accounts for 

functions of the geosynthetic reinforcement beyond simple tensile inclusions and takes into 

account the backfill and reinforcement spacing comprising the GRS composite material.(1)In this 

model, connection, reinforcement pullout, and other mechanisms associated with a tie-back 

design model are not modes of failures for closely spaced GRS.(10) Adding to the understanding 

of the composite behavior, conducting PTs without a CMU facing allowed for the failure surface 

for each GRS composite to be visually observed.

For the PTs with reinforcement spacing less than 12 inches, a shear surface formed through the 

composite (figure 34, figure 35, and figure 36), leading to rupture of the reinforcement. The 

same mode of failure was not seen for the 151⁄4-inch spaced system (figure 37) where failure was 

determined by the soil between the sheets of geotextile, and the full strength of the reinforcement 

was never developed.
39

Figure 33. Graph. Repeatability of PT at TFHRC.
40

Figure 34. Photo. TF-11 at failure with Sv = 3-13/16 inches, Tf = 1,400 lb/ft, 

and 21A material.

Figure 35. Photo. TF-3 at failure with Sv = 75⁄8 inches, Tf = 2,400 lb/ft, and 21A material.
41

Figure 36. Photo. TF-13 at failure with Sv = 111⁄4 inches, Tf = 3,600 lb/ft, and 21A material.

Figure 37. Photo. TF-10 at failure with Sv = 151⁄4 inches, Tf = 4,800 lb/ft, and 21A material.
42

The same diagonal shear failure mechanism could also be seen through forensic examination of 

the geotextile after testing the PTs with facing elements. For example, figure 38 shows the 

rupture pattern in each layer of reinforcement for TF-6 (qult,emp = 43,828 psf); the lowest layer of 

reinforcement (near the base of the GRS composite) is in the foreground of the figure with 

subsequent layers laid out above. A distinct rupture pattern follows a diagonal path through the

depth of the specimen; rupture around the connections is also seen in the middle layers.

Rupture at the connections occurred due to extreme downdrag of the reinforced fill behind the 

block whereby the differential settlement between the fill and the block caused tension (and 

ultimately rupture) in the reinforcement at the connection. For example, at failure, TF-6 

experienced a total maximum vertical strain of 15.7 percent, equating to about 1 ft of settlement; 

considering the reinforcement layers are spaced every 0.64 ft, the total settlement extended 

below one layer of the GRS composite. Note that even with the rupture at the connections, only 

some cracking of the facing blocks was observed (figure 39).
43

Figure 38. Photo. Rupture pattern for geotextiles in TF-6 (qult,emp = 43,828 psf); the lowest 

layer of reinforcement is the closet fabric in the picture.
44

Figure 39. Photo. Post-test picture of TF-6 (Sv = 75⁄8 inches, Tf = 4,800 lb/ft).
45

5. COMPARISON TO PLANE STRAIN CONDITIONS

Both vertical capacity and stiffness are impacted by the boundary effects related to testing. 

Considering the application of PTs to bridge abutments, the relationship between the more plane 

stress conditions of a columnar PT and the plane strain (PS) conditions of an abutment was 

investigated.

5.1 CAPACITY

A PT with no facing can be considered a type of unconfined compression test; its state of stress 

at failure can be represented by the Mohr-Coulomb equation (figure 40).

Figure 40. Equation. Mohr-Coulomb shear strength.

Where = shear strength, = applied normal stress, c = cohesion and = friction angle. Note 

that when applying figure 40 to a GRS composite, c and represent the shear strength parameters 

of the GRS composite (cGRS and respectively) and not of the backfill soil alone. In an 

unconfined compression PT, where the facing element has been removed, the ultimate capacity 

of the GRS column (qult,PT) can be approximated using figure 41.

Figure 41. Equation. Ultimate capacity of an unconfined GRS PT.

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 footing bearing on top of a slope 

(figure 42):(24)

Figure 42. Equation. Ultimate capacity of a strip footing on slope.

Where qult,PS = ultimate capacity of strip footing under PS conditions, cGRS is the cohesion of the 

GRS composite, = unit weight of the GRS composite, b = footing width, and Ncq and Nq = 

bearing capacity factors for a strip footing with a rough base (given by Meyerhof, 1957). The 

bearing capacity factor Nq approaches zero when the slope angle ( ) is 90; thus, figure 42

reduces to figure 43 for a vertical GRS abutment.

Figure 43. Equation. Ultimate capacity of a strip footing on a vertical GRS abutment.

τ = c + σtanφ 

 σ φ

φ

φGRS qult,PT = 2cGRS

qult,PS = cGRSNcq + 0. 5γGRSbNγq

GRS βs

qult,PS = cGRSNcq
46

The ratio of the bearing capacity of a strip footing on top of a GRS abutment to that of a GRS 

column is therefore calculated according to figure 44.

Figure 44. Equation. Ratio of plane strain capacity to PT capacity.

For a surface footing (i.e., no embedment depth) 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), mathematically shown in figure 45 and graphically shown in 

figure 46. To use figure 46, the effective cohesion of the GRS composite is required to calculate 

the stability factor, Ns; this can be obtained from laboratory or numerical experiments.

Figure 45. Equation. Stability Factor.

Figure 46. Graph. Ratio of plane strain capacity to PT capacity 

for different stability factors.

Pham (2009) conducted five plane strain generic soil-geosynthetic composite (GSGC) tests: two 

of the tests were identical except one was unconfined (0 psi) and the other test had a uniform 

confining stress of 5 psi.(25) The GRS composite consisted of a granular A-1-a backfill (soil 

qult,PS

qult,PT

=

Ncq

2

Ns =

γGRSH

cGRS

0

0.5

1

1.5

2

2.5

3

0 1 2 3 4 5

qult,PS/qult,PT

a/b (for Ns = 0) or a/H (for Ns > 0)

Ns = 0

Ns = 2

Ns = 4
47

classification = GW-GM, unit weight, = 153.7 lb/ft3, cohesion, c = 1,480 psf and friction 

angle, = 50o) and a 4,800 lb/ft geotextile reinforcement spaced at 75⁄8 inches. The corresponding 

failure stresses were about 42,450 psf and 70,957 psf for the 0 and 5 psi confining stresses, 

respectively. These failure stresses are calculated as the ratio of applied vertical load to the area 

of the top load pad (i.e. footing) on the GSGC mass for each test. Note that Pham (2009) 

reported failure stresses as the ratio of applied vertical load to the total area of the GSGC 

mass.(25)

The results of these tests can be used to determine the strength properties (cohesion and friction 

angle) of a particular GRS composite. The Mohr-Coulomb failure circles for the two Pham 

(2009) tests are shown in figure 47; the resulting cohesion and friction angle for the GRS 

composite tested is 3,342 psf and 72o, respectively; the stability factor (figure 45) is 0.29. Based 

on this stability factor, the ratio of plane strain capacity to column (PT) capacity varies with a/H

(figure 48). The correction factor to convert from the PT to a plane strain condition for a typical 

GRS abutment with the setback (a) equal to 8 inches and height from 10 to 30 ft (i.e., a/H = 0.02 

to 0.07) ranges between 0.92 and 0.97 for this particular GRS composite, close to unity. 

Therefore, the failure of the PT is approximately representative of an in-service PS condition for 

well-graded gravels, meaning the PT results can be directly used to estimate the strength limit for 

capacity in design. Note that the strength properties of the foundation soil underneath in-service

GRS abutments would still need to be verified separately.

(1)

Note that there are some limitations to the formulation above. The relationship between the 

bearing capacity of the strip footing and the performance tests requires the cohesion and friction 

angle of the GRS composite, which are not easy parameters to determine. Additional work is 

required to develop relationships between the composite shear strength parameters and the soil 

shear strength parameters, with respect to reinforcement spacing and strengths.

Figure 47. Graph. Mohr-Coulomb failure envelope for Pham (2009) 

plane strain GSGC tests.

γ

φ

0

10000

20000

30000

40000

50000

60000

70000

80000

0 20000 40000 60000 80000

Shear Stress (psf)

Normal Stress (psf)
48

Figure 48. Graph. Plane strain capacity to PT capacity for a stability factor of 0.29. 

5.2 STIFFNESS 

The GRS PTs have a height to width ratio of about 2 (table 7); the stiffness of the unconfined 

GRS column (SPT) can therefore be taken as equal to the Young’s modulus of the GRS 

composite (EGRS). In contrast, if a performance test was conducted on an infinitely long 

unconfined GRS abutment (i.e., plane strain conditions), having a finite width as shown in 

figure 49, the stiffness of this GRS mass is represented according to figure 50. 

Figure 49. Illustration. Infinitely Long Unconfined GRS abutment. 

0

0.5

1

1.5

2

2.5

3

0 1 2 3 4 5

qult,PS/qult,PT

a/H (Ns = 0.29)
49

Figure 50. Equation. Stiffness of an Infinitely Long Unconfined GRS abutment. 

Where q = applied stress, = vertical strain, EGRS = Young’s modulus of the GRS and = 

Poisson’s ratio of the GRS. Assuming is equal to 0.5, this infinitely long GRS strip will be 

1.33 times stiffer than a square GRS column.(2) In reality, the width is not finite in the direction 

perpendicular to a GRS abutment, as shown on the left of figure 51. To date, typical L/b ratios 

for in-service GRS abutments range between 7 and 14, depending on load requirements, where L

is the length and b is the width of the bearing area. NAVFAC (1986) considers L/b ratios > 5 to 

be plane strain while current FHWA practice considers L/b ratios > 10 to be plane strain 

conditions; therefore, most GRS abutments can be idealized as plane strain.(26)

In the context of elastic theory, the bridge footing on a GRS abutment can be considered a strip 

running parallel to the edge of the wall at an offset, as shown on the left side of figure 51. 

Considering the response at low strain levels is close to elastic conditions, the stiffness of such a 

strip surface footing can be estimated using elastic theory, in particular the solution for a strip 

footing on a linear elastic half space.(27,28) Since the conditions are not a linear elastic half space 

for abutment applications, the solution for a strip footing on top of a wall is obtained by doubling 

the solution for a strip footing on a linear elastic half space, as shown on the right side of 

figure 51. Note that this assumes zero lateral deformations along the wall height, which is not 

strictly correct for GRS abutments since the CMU facing is free to displace laterally; however, it 

is the best available method to date to estimate vertical settlement. 

Figure 51. Illustration. Solution for strip footing on top of a wall. 

For a material with a Poisson’s ratio of 0.5, Gibson (1967) provided solutions for the relative 

displacement between the center of the strip footing and any arbitrary point assuming a constant 

stiffness with depth.(27) If the arbitrary point is selected to be the bottom of the GRS wall, the 

expression for settlement without doubling the load is shown in figure 52. 

q

εv

=

EGRS

(1 + VGRS)(1 − VGRS)

v GRS GRS
50

Figure 52. Equation. Vertical displacement of a GRS abutment with a strip footing. 

Where q = applied stress, b is the footing width, EGRS is the Young’ modulus of the GRS 

composite, a is the setback distance between the face of the wall and the applied load, and H is 

the height of the abutment. The corresponding vertical strain ( ) is given by figure 53. 

Figure 53. Equation. Vertical strain. 

Where is the vertical displacement (figure 52) and H is the height of the wall. 

The stiffness of a GRS abutment supporting a surface strip footing (SGRS) can be related to the 

applied stress (2q, due to doubling of the load) and the vertical strain ( ) (figure 54).

Figure 54. Equation. Stiffness of a GRS abutment supporting a strip footing. 

The ratio between this stiffness and the stiffness from a performance test (SGRS = EGRS) can then 

be calculated according to figure 55. The variation of stiffness ratio with b/H is plotted in 

figure 56. It was found that this ratio is essentially independent of the value of a/b.

Figure 55. Equation. Vertical displacement of a GRS abutment with a strip footing. 

ρ =

3qb

4πEGRS 

1

2

 1 +

a +

b

2

b

2

 ln 

H2 + (a +

b

2

)

2

(

b

2

)

2

 +

1

2

 1 −

a +

b

2

b

2

 ln 

H2 + a

2

(

b

2

)

2

 

+

H

a

tan−1 

a + b

H

 +

H

a

tan−1 

−b

H

 

v

εv =

ρ

H

ρ

v

SGRS =

2q

εv

=

2qH

ρ

SGRS

SPT

=

2qH

ρEGRS

=

8πH

3b 

1

2

 1 +

a +

b

2

b

2

 ln 

H2 + (a +

b

2

)

2

(

b

2

)

2

 +

1

2

 1 −

a +

b

2

b

2

 ln 

H2 + a2

(

b

2

)

2

 +

H

a

tan−1 

a + b

H

 +

H

a

tan−1 

−b

H

 
51

Figure 56. Graph. Ratio of plane strain stiffness of a strip footing on top of a wall (SGRS) to 

that of a PT (SPT) for the case of constant stiffness with depth.

For GRS abutments, b is a minimum of 2.5 ft for bridge spans larger than 25 ft and has been as 

large as 4.5 ft for the Tiffin River Bridge; H typically ranges from 10 to 30 ft; b/H is therefore 

between about 0.08 and 0.45.(5)From figure 56, the ratio of plane strain stiffness of a strip 

footing on top of a wall (SGRS) to that of a PT (SPT) ranges between 0.92 and 3 suggesting that the 

stiffness reported from a PT will be conservative (up to a factor of three times) compared to what 

a GRS abutment might experience in-service.

Note that further research should be conducted on the stiffness of GRS abutments, and the 

corresponding equivalency to a GRS PT, since the above formulations are simplistic and do not 

consider the effects of the facing element (e.g., CMUs) on the capacity and stiffness of a GRS 

composite. The effect of compaction-induced stresses and related shakedown is also ignored. 

Compaction-induced stresses increase the locked-in lateral stress especially in the upper reaches 

of the GRS; associated with an increase in confinement is an increase in stiffness and its effects 

should be investigated.

0

0.5

1

1.5

2

2.5

3

3.5

4

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

SGRS / SPT

b/H

53

6. PARAMETRIC ANALYSIS

The combined series of tests performed in Defiance County and at TFHRC make up a parametric 

study where the following parameters were investigated: (1) aggregate type, (2) compaction, 

(3) gradation, (4) bearing bed reinforcement, (5) reinforcement strength, (6) similar ratios of 

reinforcement strength to spacing, (Tf/Sv), and (7) facing.

6.1 EFFECT OF AGGREGATE TYPE

The aggregate selected has a large impact on the performance and composite behavior of GRS; 

both the size (i.e., dmax) and the strength (i.e., c, ) of the backfill selected are important. In PTs, 

it is very difficult to isolate one backfill variable over another to determine the individual 

contributions to GRS performance. Instead, general conclusions based on aggregate 

classification and angularity can be drawn using PT results; however, advanced computer 

modeling is a potential avenue to better understand the component contributions of particular fill 

materials.

The majority of GRS- Integrated bridge systems (IBSs) built across the country use an opengraded aggregate for the abutment backfill. Commonly, an AASHTO No. 89 has been employed; 

an early performance test performed in Defiance County shows the behavior of a GRS composite 

with this backfill and 4,800 lb/ft reinforcement spaced at 75⁄8 inches.

(1) To create a database of 

PT results with alternate aggregate types, while providing an opportunity to demonstrate the 

construction of a PT to FHWA Division Offices, Federal Lands Highway Division, and the 

Resource Center through the EDC GRS-IBS Validation Sessions, a series of tests were 

conducted in this study (table 9). Note that since these PTs were conducted as demonstration 

piers, the level of compaction control on each lift was not consistent or uniform.

Table 9. Parametric study on aggregate size.

Test

No.

Backfill Reinforcement

Type (°) Facing

c

(psf)

dmax

(inch)

Tf

(lb/ft)

Sv

(in.)

DC-1 8 54 0 1⁄2 4,800 75⁄8** CMU

DC-2 8P* 46 0 3⁄4 4,800 75⁄8** CMU

DC-3 57 52 0 1 4,800 75⁄8** CMU

DC-4 9 49 0 3⁄8 4,800 75⁄8** CMU

φ = the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, and Sv = the spacing.

*Rounded pea-gravel angularity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

Based on the results (table 10), no relationship can be determined with respect to the strength 

limit for open-graded aggregates since there are only two PTs conducted to failure of the GRS 

composite. The largest aggregate tested, the No. 57 stone, had the lowest service limit of all the 

tests, indicating more deformation under an applied load. In addition, the rounded pea gravel had 

a lower strength and service limit than the more angular aggregate meeting the same gradation 

specifications for an AASHTO No. 8 material.

φ

φ
54

Table 10. Effect of aggregate type results.

Maximum Tested

Strength 

Limit Design Limit

Service 

Limit

Test

No.

qmax

(psf) (%)

qult,emp

(psf) (psf)

Vallow = 

/3.5 (psf) (psf)

DC-1 23,310 7.95 23,310 19,983 5,709 3,065

DC-2 22,709 7.07 22,709 19,399 5,543 2,171

DC-3 18,447 5.82 N/A 16,182 4,623 1,324

DC-4 26,730 7.64 N/A 17,350 4,957 2,212

qmax = the maximum applied pressure during testing, = the maximum recorded 

vertical strain, qult,emp = the measured failure pressure, = the applied stress at 

5-percent vertical strain, Vallow,emp = the total allowable pressure on the GRS,

(1)

= the applied stress at 0.5-percent vertical strain.

6.2 EFFECT OF COMPACTION

The effect of compaction for open-graded materials was investigated through two PTs conducted 

in Defiance County, OH (table 11). The testing conditions were identical except for the 

compaction effort. DC-1 was compacted to non-movement while DC-5 involved no compaction 

effort beyond end dumping the material and leveling for each lift of fill. The results show a 

similar vertical capacity (23,310 psf versus 21,539 psf for DC-1 and DC-5, respectively); 

however, the higher strength for DC-1 is likely due to the increased locked-in stresses induced 

due to compaction. Note that the friction angle of both composites is reported as the same; 

however, the degree of compaction may have a small impact for open-graded aggregates (larger 

for well-graded backfills). This effect is currently being investigated.

While the capacity of the specimens was similar, the deformation response was different

(figure 57). As expected, Test DC-5 had a softer response resulting from not being compacted. 

The modulus of the primary settlement portion of the curve for the uncompacted composite 

(DC-5) is about 270 ksf while the modulus for the compacted composite (DC-1) is 430 ksf 

(table 11).

εmax q@ε≤5% q@ε≤5% q@ε=0.5% εmax q@ε=5% q@ε=0.5%
55

Table 11. Parametric study on compaction.

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

Tf

(lb/ft)

Sv

(in.)

DC-1 8 54 0 4,800 75⁄8** CMU 430 23,310

DC-5 8*** 54 0 4,800 75⁄8** CMU 270 21,539

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

***Uncompacted sample.

At the service limit state of 0.5 percent vertical strain, the allowable stress is limited to 316 psf 

for the uncompacted GRS composite, but 3,065 psf for the same composite compacted to nonmovement (table 8); however, if negating the initial immediate settlement resulting from the first 

load increment, before the onset of primary deformation of the uncompacted sample, the 

allowable stress is about 1,345 psf (figure 58).

The ratio of strain for the uncompacted (DC-5) and compacted (DC-1) GRS composite 

( / ), accounting for the immediate deformation related to the uncompacted 

composite (DC-5), decreases with increasing applied pressure (figure 59). For the particular GRS 

composite tested (table 11, figure 58), at an in-service dead load of 4,000 psf, an uncompacted 

abutment will experience about 3 times the strain as a compacted abutment; however, near 

failure, at about 20,000 psf, an uncompacted abutment will experience about 1.7 times the strain 

as a compacted abutment. Note that in the evaluation of tolerable settlements for highway bridge 

design, FHWA recommends the construction-point concept, whereby the settlements between

critical construction points (such as between application of dead load and opening to traffic) are 

evaluated.

(29)

φ

φ

εv,uncompact εv,compact
56

Figure 57. Graph. Comparison between compacted and uncompacted GRS composites.
57

Figure 58. Design service limit for uncompacted sample DC-5.
58

Figure 59. Graph. Comparison of compacted and uncompacted strains 

between the DC-1 and DC-5 tests.

6.3 EFFECT OF BEARING BED REINFORCEMENT

Bearing bed reinforcement, where the reinforcement is spaced at half the primary spacing, is 

recommended in at least the top five courses of CMU facing elements for GRS abutments to aid 

in serviceability.

(1) To investigate the impact of the bearing bed, two PTs were conducted with 

identical parameters, except one (TF-8) had two courses of bearing bed reinforcement, as 

recommended in the empirical design procedure using performance testing by Adams et al. 

2011a (table 12); the other (TF-7) had no bearing bed reinforcement.

(1) The axial behavior results 

indicate that the bearing bed provides some added vertical capacity; however, vertical 

deformation is not improved at low strain levels (figure 60). The modulus for primary 

compression is similar whether or not the bearing bed is present (table 12).
59

Table 12. Parametric study on bearing bed reinforcement.

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(in.)

Tf

(lb/ft)

Sv

(in.)

TF-7 21A 53 115 1 4,800 75⁄8 no CMU 320 26,546

TF-8 21A 53 115 1 4,800 75⁄8** no CMU 320 29,134

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

Figure 60. Graph. Effect of bearing bed reinforcement for TF-7 and TF-8.

Looking at the lateral deformation characteristics for both tests at an applied stress of about 

3,600 psf, typical of bridge loads (figure 61), the bearing bed reinforcement serves to limit 

deformation in the zone of its placement. The approximate location of the bearing bed 

reinforcement is illustrated with red dashed lines in figure 61. At considerably higher loads,

around 26,600 psf, the same effect is observed (figure 62). Note that the LVDT located at the 

bottom of the PT for the TF-7 test was damaged during testing due to sloughing of the fill 

φ

φ
60

material so no values were recorded in the later stages of the test; the full lateral displacement 

curves for all of the PTs will be presented in a separate report.

At service loads, when bearing bed reinforcement is present (TF-8), the maximum lateral 

deformation is lower and occurs at about mid-height (381⁄8 inches from the top) whereas when 

bearing bed reinforcement is excluded (TF-7), the maximum lateral deformation occurs near the 

top (75⁄8 inches from the top). The lateral deformation at 75⁄8 inches below the top of the GRS PT 

is 0.16 inches for TF-7 (no bearing bed) and 0.07 inches for TF-8 (with bearing bed 

reinforcement); the bearing bed reduced the lateral deformation by about half at 3,600 psf 

applied vertical pressure. At larger loads, the difference in the measured lateral deformation

between including a bearing bed and not diminishes considerably at the top of wall. Note that 

since the results are based on only two PTs, additional testing is required to verify the 

conclusions for different aggregate and reinforcement materials and to investigate the depth of

influence for the bearing bed.

Figure 61. Graph. Measured lateral deformation at 3,600 psf applied stress for TF-7 (no 

bearing bed reinforcement) and TF-8 (2 courses of bearing bed reinforcement).
61

Figure 62. Graph. Measured lateral deformation at 26,600 psf applied stress for TF-7 (no 

bearing bed reinforcement) and TF-8 (2 courses of bearing bed reinforcement).

6.4 Effect of Gradation

For the GRS-IBS, both open-graded and well-graded materials that meet the specifications in the 

Interim Implementation Guide are acceptable for use in the GRS abutment; however, there are 

advantages and disadvantages of both.(1) Open-graded materials are free draining, easier to work 

with, and their use in construction is independent of weather conditions; however, they are less 

stiff than well-graded materials, which can achieve greater density under the same compactive 

effort. In addition, the classical bell-shaped Proctor curves cannot be attained with open-graded

materials. The primary disadvantage of working with well-graded fill is compaction control and 

maintaining the optimum moisture content for efficient compaction.

PTs TF-1 and TF-2 are identical except for the backfill material type; TF-1 used an open-graded

AASHTO No. 8 while TF-2 used a well-graded AASHTO A-1-a backfill (table 13). The friction 

angle for both tests were similar (55° versus 53° for TF-1 and TF-2, respectively), but the wellgraded material (TF-2) had cohesion of 115 psf. The corresponding soil shear strengths at a 

given applied stress (figure 40) are therefore only 5 percent different, with TF-2 having slightly 

higher shear strength.
62

Table 13. Parametric study on gradation (Tf = 2,400 lb/ft, Sv = 75⁄8 inches).

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(inch)

Tf

(lb/ft)

Sv

(inch)

TF-1

++ 8 55 0 1⁄2 2,400 75⁄8 CMU 320 20,487

TF-2 21A 53 115 1 2,400 75⁄8 CMU 710 25,260

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

++Technical difficulties resulted in unloading/reloading of the composite.

The resulting load-deformation profiles are shown in figure 63. The well-graded material is 

considerably stiffer than the open-graded material. At an applied pressure of 4,000 psf, the 

vertical strain is about 1.1 percent for TF-1 (open-graded) and 0.4 percent for TF-2 (wellgraded). The modulus of the composite tested in TF-1 is 320 ksf (table 13) compared to 710 ksf 

for the composite tested in TF-2; TF-1 (open-graded) is 55 percent less stiff than TF-2 (wellgraded). In terms of bearing capacity, TF-1 was 20 percent less strong than TF-2. The results 

indicate that the gradation, and perhaps cohesion, impacts the stiffness more so than strength,

indicating that well-graded fills have an advantage with respect to serviceability. As previously 

discussed in section 6.1, isolating the effect of cohesion on the performance is difficult using 

PTs, although based on the soil-geosynthetic capacity equation (figure 8), it will serve to 

improve capacity, although its contribution should not be considered in design.

φ

φ
63

Figure 63. Graph. Comparison of open-graded and well-graded backfills

for TF-1 and TF-2.

Comparing DC-1 and TF-6 also provides similar insight into the modulus difference between 

open-graded and well-graded composites, respectively (table 14). Note that DC-1 included two

courses of bearing bed reinforcement; however, it was previously shown that the bearing bed 

reinforcement does not impact the modulus at low strain levels, but the capacity is slightly 

improved with the two additional layers of reinforcement at the top. The modulus of the 

composite with open-graded backfill (DC-1) was 430 ksf, whereas the modulus of the composite 

with well-graded backfill (TF-6) was 750 ksf. Both are slightly larger than the measured 

modulus of similar composites with lower reinforcement strengths (table 13).
64

Table 14. Parametric study on gradation (Tf = 4,800 lb/ft, Sv = 75⁄8 inches.)

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(inch)

Tf

(lb/ft)

Sv

(inch)

DC-1 8 54 0 1⁄2 4,800 75⁄8** CMU 430 23,310

TF-6

++ 21A 53 115 1 4,800 75⁄8 CMU 750 43,763

φ

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

++Technical difficulties resulted in unloading/reloading of the composite.

6.5 EFFECT OF REINFORCEMENT STRENGTH

The impact of reinforcement strength on the behavior of a GRS composite was investigated for 

both open-graded (table 15) and well-graded aggregates (table 16).

Table 15. Parametric study on reinforcement strength with open-graded aggregates.

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(inch)

Tf

(lb/ft)

Sv

(inch)

DC-1 8 54 0 1⁄2 4,800 75⁄8** CMU 430 23,310

TF-1

++ 8 55 0 1⁄2 2,400 75⁄8 CMU 320 20,487

φ

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

**Two courses of bearing bed reinforcement placed at the top of the PT.

++Technical difficulties resulted in unloading/reloading of the composite.

Table 16. Parametric study on reinforcement strength with well-graded aggregates.

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(inch)

Tf

(lb/ft)

Sv

(inch)

TF-2 21A 53 115 1 2,400 75⁄8 CMU 710 25,260

TF-6

++ 21A 53 115 1 4,800 75⁄8 CMU 750 43,763

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

++Technical difficulties resulted in unloading/reloading of the composite.

DC-1 and TF-1 both tested similar open-graded aggregates (AASTHO No. 8s) at the same 

reinforcement spacing of 75⁄8 inches, but DC-1 used a 4,800 lb/ft geotextile while TF-1 used a 

2,400 lb/ft geotextile (table 15). The response indicates that the higher reinforcement strength 

φ

φ

φ

φ
65

(4,800 lb/ft) produces a stiffer and stronger response than the lower reinforcement strength 

(2,400 lb/ft) for open-graded backfill. By doubling the reinforcement strength, the results 

indicate an increase by a factor of 1.14 for capacity and 1.34 for the initial stress-strain ratio.

Comparing the effect of reinforcement strength with a well-graded AASHTO A-1-a aggregate, 

with and without facing, it was found that increasing the ultimate reinforcement strength

(MARV) by a factor of two from 2,400 lb/ft (TF-2) to 4,800 lb/ft (TF-6) at the same 

reinforcement spacing (Sv = 75⁄8 inches) results in an increase in the measured capacity and initial 

stress-strain ratio by a factor of 1.73 and 1.06, respectively for CMU facing (table 16). When no 

facing is present, the same increase in reinforcement strength from 2,400 lb/ft (TF-3) to 

4,800 lb/ft (TF-7) at a spacing of 75⁄8 inches results in an increase in the measured capacity by a 

factor of 1.5. For open-graded aggregates, the percent increase in capacity is less than for the 

well-graded material tested which may be due to the increased cohesion and maximum density of 

the well-graded material, thus leading to higher soil shear strengths and increased stiffness 

properties.

6.6 EFFECT OF THE RELATIONSHIP BETWEEN REINFORCEMENT STRENGTH 

AND SPACING

The current AASHTO LRFD Bridge Design Specifications (2012) do not distinguish between 

MSE technology and closely-spaced GRS technology.

(30)

In AASHTO (2012), reinforcement

spacing is linearly proportional to the reinforcement strength, leading engineers to prefer larger 

spaced systems with proportionally greater reinforcement strengths.(10) Research on closely 

spaced systems, however, indicates that reinforcement spacing plays a significantly larger role 

than the reinforcement strength.

(5,11 ,25 ,31)

To investigate the relationship between the MARV wide width reinforcement tensile strength 

(Tf) and spacing (Sv), several PTs were designed to keep the same Tf /Sv ratios, both with 

(table 17) and without facing elements (table 18). The strength of the reinforcement was assumed 

to be the manufacturer supplied MARV (table 6). Within the data set of this study, there are three 

pairs of tests (with and without facing) at a constant Tf /Sv ratio (table 19). For a Tf /Sv ratio of 

3,800 lb/ft2, the load-test curves for the tests with facing are presented in figure 64 and the loadtest curves for the tests without facing are presented in figure 65.

Table 17. Parametric study for 3,800 lb/ft2 Tf/Sv ratio (with facing).

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(in.)

Tf

(lb/ft)

Sv

(in.)

TF-2 21A 53 115 1 2,400 75⁄8 CMU 710 25,260

TF-9 21A 53 115 1 4,800 151⁄4 CMU 550 22,310

TF-14 21A 53 115 1 3,600 111⁄4 CMU 460 23,562

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

φ

φ
66

Table 18. Parametric study for 3,800 lb/ft2 Tf/Sv ratio (with no facing).

Test

No.

Backfill Reinforcement

Facing

Eo

(ksf)

qult,emp

Type (°) (psf)

c

(psf)

dmax

(in.)

Tf

(lb/ft)

Sv

(in.)

TF-3 21A 53 115 1 2,400 75⁄8 no CMU 330 17,491

TF-10 21A 53 115 1 4,800 151⁄4 no CMU 260 10,330

TF-13 21A 53 115 1 3,600 111⁄4 no CMU 260 12,960

= the peak friction angle, c = the cohesion, Tf = the MARV value of the wide width tensile 

strength, Sv = the spacing, Eo = the initial stress-strain ratio, and qult,emp = the measured vertical 

capacity.

Table 19. Tf/Sv ratios for each PT.

No.

Tf/Sv

3,800 lb/ft24,400 lb/ft27,600 lb/ft2

CMU Facing No Facing CMU Facing No Facing CMU Facing No 

Facing

1 TF-2 TF-3 TF-12 TF-11 TF-6TF-7

2 TF-9 TF-10 TF-8

3 TF-14 TF-13 DC-1

4 TF-1 DC-2

5 DC-3

6 DC-4

7 DC-5

Note: The table shows which sets of tests were performed at a particular Tf/Sv ratio, thus some 

cells are blank.

φ

φ
67

Figure 64. Graph. Stress-strain curves for PTs with CMUs at Tf/Sv = 3,800 psf
68

Figure 65. Graph. Stress-strain curves for PTs with no CMU facing at Tf/Sv = 3,800 psf.
69

For a Tf/Sv ratio of 3,800 lb/ft2, as the reinforcement spacing increased, the vertical capacity 

decreased for the same Tf/Sv ratio, whether a CMU facing was absent (figure 66) or present 

(figure 67). Contrary to MSE design theory, as reinforcement strength increased, while

increasing the spacing proportionally, the vertical capacity decreased, whether a CMU facing 

was absent (figure 68) or present (figure 69).

This suggests that the relationship between reinforcement strength and spacing is not 

proportional to capacity as outlined in current MSE design; a GRS abutment with a given Tf and 

Sv will not have the same strength figure as a GRS abutment with twice the strength (2Tf) and 

reinforcement spacing (2Sv).

Figure 66. Graph. Capacity of GRS with no CMU facing at various reinforcement spacing 

for different Tf/Sv ratios.

0

5000

10000

15000

20000

25000

30000

0 5 10 15 20

Vertical Capacity (psf)

Reinforcement Spacing (inch)

Tf/Sv = 3800 psf

Tf/Sv = 4400 psf

Tf/Sv = 7600 psf

No CMU Facing
70

Figure 67. Graph. Capacity of GRS with CMU facing at various reinforcement spacing for 

different Tf/Sv ratios.

Figure 68. Graph. Capacity of GRS with no CMU facing at various reinforcement strength 

for different Tf/Sv ratios.

0

5000

10000

15000

20000

25000

30000

35000

40000

45000

50000

0 5 10 15 20

Vertical Capacity (psf)

Reinforcement Spacing (inch)

Tf/Sv = 3800 psf

Tf/Sv = 4400 psf

Tf/Sv = 7600 psf

CMU Facing

0

5000

10000

15000

20000

25000

30000

0 2000 4000 6000

Vertical Capacity (psf)

Reinforcement Strength (lb/ft)

Tf/Sv = 3800 psf

Tf/Sv = 4400 psf

Tf/Sv = 7600 psf

No CMU Facing
71

Figure 69. Graph. Capacity of GRS with CMU facing at various reinforcement strength for 

different Tf/Sv ratios.

At the same Tf/Sv ratio of 3,800 lb/ft2, increasing the reinforcement spacing and reinforcement 

strength by a factor of two from 75⁄8 to 151⁄4 inches and 2,400 to 4,800 lb/ft, respectively, resulted 

in a reduction of the capacity by a factor of 0.9 and 0.6 for CMU facing and no facing, 

respectively. According to the design theory employed by AASHTO (2012), there would be no 

reduction in capacity.(10) The relationship between reinforcement strength and spacing is 

therefore not directly proportional. The results are similar to those reported by Pham (2009).(25)

6.7 EFFECT OF FACING

In the GRS-IBS Interim Implementation Guide, the effect of the facing is ignored in determining 

the capacity of a GRS composite (i.e., confining stress is equal to zero in figure 8).

(1) The PTs 

provide insight on the magnitude of the impact CMU facing elements have on the performance 

of GRS composites. Of the nineteen tests included in this study, there were five pairs of identical 

GRS composites constructed with the well-graded aggregate; each pair consisted of one test with 

a frictionally connected CMU facing and another test without the CMU facing (table 21). The 

stress-strain response for each pair is shown in figure 70 through figure 74.

0

5000

10000

15000

20000

25000

30000

35000

40000

45000

50000

0 2000 4000 6000

Vertical Capacity (psf)

Reinforcement Strength (lb/ft)

Tf/Sv = 3800 psf

Tf/Sv = 4400 psf

Tf/Sv = 7600 psf

CMU Facing
72

Figure 70. Graph. Stress-strain response for TF-2 (CMU facing) and TF-3 (no CMU 

facing) with Sv = 75⁄8 inches and Tf = 2,400 lb/ft.

Figure 71. Stress-strain response for TF-6 (CMU facing) and TF-7 (no CMU facing) with 

Sv = 75⁄8 inches and Tf = 4,800 lb/ft.
73

Figure 72. Graph. Stress-strain response for TF-9 (CMU facing) and TF-10 (No CMU 

facing) with Sv = 151⁄4 inches and Tf = 4,800 lb/ft.
74

Figure 73. Graph. Stress-strain Response for TF-12 (CMU facing) and TF-11 (no CMU 

facing) with Sv = 3-13/16 inches and Tf = 1,400 lb/ft.

Figure 74. Graph. Stress-strain response for TF-14 (CMU facing) and TF-13 (no CMU 

facing) with Sv = 111⁄4 inches and Tf = 3,600 lb/ft.
75

The frictionally connected CMU facing has an impact on the performance of GRS; it provides 

confinement, leading to a stiffer response and an increased capacity compared to a GRS 

composite with no facing element (table 20). In addition, at the current service stress limit 

(applied stress, q, of 4,000 psf), the ratio of service vertical strain with no facing to service 

vertical strain with a CMU facing ranges from 1.2 to 2.2 (table 21). The largest impact was for 

the 155⁄8-inch spaced GRS composite (TF-9 and TF-10). At ultimate failure, the ratio is 

considerably closer, ranging from 0.8 to 1.2.

Table 20. Effect of CMU facing on stiffness and capacity.

Test Facing

Sv

(inch)

Tf/Sv

(psf)

Eo

(ksf)

Eo,CMU

Eo,no 

CMU

qult,emp

(psf)

qult, emp CMU

qult, emp no CMU

TF-2 CMU 75⁄8 3,800

710

2.15 25,260 1.44

TF-3 None 330 17,491

TF-6 CMU 75⁄8 7,600

750

2.34 43,763 1.65

TF-7 None 320 26,546

TF-9 CMU 151⁄4 3,800

550

2.12 22,310 2.16

TF-10 None 260 10,330

TF-12 CMU 3-13/16 4,400

810

2.08 29,030 1.25

TF-11 None 390 23,249

TF-14 CMU 111⁄4 3,800

460

2.09 23,562 1.82

TF-13 None 220 12,960

Sv = the reinforcement spacing, Tf = the MARV value of the wide width tensile strength, 

Eo = the initial stress-strain ratio, Eo,CMU = the initial stress-strain ratio for tests with CMU facing, 

Eo, no CMU = the initial stress-strain ratio for tests without any facing, qult,emp = the measured 

vertical capacity, qult,emp CMU = the measured failure pressure for tests with CMU facing, and 

qult,emp no CMU = the measured failure pressure for tests without any facing.
76

Table 21. Effect of CMU facing on strain.

Test Facing

Sv

(inch)

Tf/Sv

(psf)

ε@q=4000psf

(%)

ε@q=4000psf, no CMU

ε@q=4000psf, CMU

ε@qult

(%)

ε@qult, no CMU

ε@qult, CMU

TF-2 CMU 75⁄8 3,800 0.39 1.84 11.46 1.20

TF-3 None 0.73 13.80

TF-6 CMU 75⁄8 7,600 0.55 1.86 15.70 0.80

TF-7 None 1.02 12.50

TF-9 CMU 151⁄4 3,800 0.74 2.16 15.60 0.91 TF-10 None 1.59 14.27

TF-12 CMU 3

13/16 4,400 0.50 1.59 13.37 0.96 TF-11 None 0.79 12.79

TF-14 CMU 111⁄4 3,800 0.93 1.17 12.69 0.97 TF-13 None 1.09 12.32

Sv = the reinforcement spacing, Tf = the MARV value of the wide width tensile strength, 

ε@q=4000psf = the measured strain at an applied load of 4,000 psf, and ε@qult = the measured strain 

at failure.

From table 20, the facing more than doubles the initial stress-strain ratio as compared to the PTs 

without any facing. In terms of capacity, the facing plays the biggest and smallest role for the 

largest spaced (Sv = 151⁄4 inches) and the closest spaced (Sv = 313/16 inches) system tested, 

respectively. The design assumption to not include the effect of confinement from the face in 

determining the capacity and required reinforcement strength is therefore conservative.(1)

While the magnitude of strain at failure is similar for a given GRS composite tested with or 

without a facing (table 21), the ultimate capacity is increased when a facing element is present

(table 20 and figure 75). For the five pairs of tests conducted at TF (table 21), including the 

CMU facing produced an improved ultimate capacity between 1.25 and 2.2 times greater than 

the GRS composite without any facing (a similar trend to the strain at the current 4,000 psf

service limit).

For the same Tf /Sv ratio of 3,800 lb/ft2, there is a linear relationship between the reinforcement 

spacing (Sv) and the ratio of capacity with CMU facing (qult,emp CMU) to the capacity without a 

facing (qult,emp no CMU) (figure 75). Similarly, there is a linear relationship with reinforcement 

strength (figure 76). To further investigate this, additional tests should be conducted at other 

Tf /Sv ratios (4,400 and 7,600 psf) and at larger reinforcement spacing.
77

Figure 75. Graph. Effect of CMU facing on ultimate capacity as a function of 

reinforcement spacing.

Figure 76. Graph. Effect of CMU facing on ultimate capacity as a function of 

reinforcement strength.
78

Using figure 8, the confining stress due to the facing elements can be back-calculated using the 

measured ultimate capacity from the PTs (figure 77). The results indicate that as reinforcement 

spacing increases, the effect of the facing element on the capacity is more pronounced. Note that 

the confining stress changes throughout the PT with applied pressure; figure 77 represents only 

the back-calculated confining stress at failure. 

Using the equation developed by Wu et al. (2010) (figure 11), the estimated confining stress for 

the CMU blocks is about 72 psf, lower than that estimated at failure (figure 77).

(11) The bulk unit 

weight and depth of the CMU is 150 pcf and 75⁄8 inches, respectively; the interface friction angle 

between the geotextile and the CMU block was assumed equal to 37°, based on connection 

strength testing performed at TFHRC.

(32) This simple method of determining the confining stress 

was used when comparing the measured capacity for each test to the estimated capacity for each 

test using figure 8.

Figure 77. Graph. Calculated confining pressure due to CMU facing at the ultimate 

capacity. 
79

7. APPLICATIONS OF PERFORMANCE TESTING TO DESIGN

The load test results from GRS performance testing have been used in the design of GRS 

abutments. Designers can use the empirical results to estimate capacity (strength limit) and 

deformation (service limit) of the GRS abutment, needed in the internal stability design.(1)In 

addition, the database of results can be used to validate the GRS capacity equation (figure 8) and 

calibrate a more accurate LRFD resistance factor.

7.1 DESIGN DATABASE

The results indicate that, based on the current limitations set forward by FHWA for GRS 

abutments of a 4,000-psf service pressure and a 0.5-percent vertical deformation service limit, 

seven GRS composites tested in this study may meet the strength and service limits (table 22).

(1)

The service limit pressure at a vertical strain of 0.5 percent is for the dead load only since live 

loads have not been found to impact the total strain for in-service GRS-IBS structures.

(5) Note 

that only three GRS composites tested in this study meet both the design limits and the material 

specifications (Tf ≥ 4,800 lb/ft, dmax ≥ 1⁄2 inches); however, based on the results of this study, 

consideration should perhaps be given to amend the criteria in future design guidance.

Table 22. PTs meeting GRS strength and service limit design criteria.

Test

No.

GRS Composite

Design Strength Limit Service 

ASD LRFD Limit

(°)

dmax

(inch)

Sv

(inch)

Tf

(lb/ft) (psf)

/FS

(psf)

( )

(psf) (psf)

DC-1 54 1⁄2 75⁄8 4,800 19,983 5,709 8,992 3,065

DC-2 46 3⁄4 75⁄8 4,800 19,399 5,543 8,730 2,171

DC-4 49 1 75⁄8 4,800 17,350 4,957 7,808 2,212

TF-2 53 3⁄8 75⁄8 2,400 18,711 5,346 8,420 4,759

TF-6 53 1⁄2 75⁄8 4,800 22,007 6,288 9,903 3,704

TF-12 53 1⁄2 3

13/16 1,400 18,573 5,307 8,358 4,028

TF-14 53 1⁄2 3

13/16 1,400 16,748 4,785 7,537 2,037

= the peak friction angle, dmax = the maximum aggregate size, Sv = the reinforcement 

spacing, Tf = the MARV value of the wide width tensile strength, = the stress at 

5-percent vertical strain from a corresponding PT, FS = the factor of safety (equal to 3.5), 

= the resistance factor for capacity (equal to 0.45), = the service stress at 

0.5-percent vertical strain from a corresponding PT, ASD = Allowable Stress Design, and 

LRFD = load and resistance factor design.

φ q@ε=5% q@ε=5% Φcap q@ε=5% q@ε=0.5% φ

q@ε=5% Φcap q@ε=0.5%
80

7.2 STRENGTH LIMIT 

7.2.1 Analytical 

PT results can be used to validate the ultimate capacity equation (figure 8). Figure 78 

illustrates the comparison between the measured capacity from various PTs and the 

calculated capacity from figure 8. Note that figure 78 includes results from additional sources 

of research.(1,4 ,8,11 ,33 ,34) Based on the results of the PTs presented in this report and plotting 

the bias against the standard normal variable (figure 79 and table 23), the mean bias is 0.88 (i.e., 

the bias at a standard normal variable of zero) with a coefficient of variation (COV) of about 

35 percent. Including other results from the literature, the average bias is 0.95 with a COV of 

about 32 percent (figure 80 and table 24). 

The mean bias values near unity and low COV values for all of the GRS composite tests (table 

24) indicate good agreement between the GRS capacity equation (figure 8) and the measured 

results. Note that the measured results used for comparison were taken at failure of the 

composite, not at a design limit of 5 percent vertical strain, as will be discussed later in the 

empirical method of design (see section 7.2.2). This may change the level of reliability between 

the analytical method and the empirical method of design, since both have equal factors of safety 

of 3.5.(1) A full LRFD calibration needs to be conducted to determine a more reliable resistance 

factor for each strength limit design method. 

Figure 78. Graph. Comparison of predicted capacity and measured capacity. 
81

Figure 79. Graph. Cumulative distribution function plot for DC and TF PTs.

y = 3.28x - 2.89

R2 = 0.92

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

2.5

0 0.5 1 1.5 2 Standard Normal Veriable, z

Bias, λ
82

Table 23. Predicted and measured vertical capacity for DC and TF PTs.

Test

Measured Predicted Bias, 

Measured

Predicted

Standard Normal 

Variable, z

qult,emp 

(psf)

qult,an,c

(psf)

DC-1 23,310 30,439 0.77 -0.32

DC-2 22,709 25,952 0.88 0.16

DC-3 N/A 40,812 N/A N/A

DC-4 N/A 17,078 N/A N/A

DC-5 21,539 30,439 0.71 -0.49

TF-1 20,487 16,447 1.25 1.15

TF-2 25,260 23,403 1.08 0.67

TF-3 17,491 22,741 0.77 -0.16

TF-4 N/A 44,786 N/A N/A

TF-5 25,920 44,786 0.58 -1.56

TF-6 43,763 45,448 0.96 0.38

TF-7 26,546 44,786 0.59 -1.19

TF-8 29,134 44,786 0.65 -0.93

TF-9 22,310 15,369 1.45 1.56

TF-10 10,330 14,707 0.70 -0.54

TF-11 23,249 28,348 0.82 0.07

TF-12 29,030 29,001 1.00 0.54

TF-13 12,960 18,764 0.69 -0.72

TF-14 23,562 19,426 1.21 0.93

Mean Bias 0.88

Standard Deviation 0.30

Coefficient of Variation (percent) 0.35

N/A = Not applicable, did not fail the composite.

λ
83

Figure 80. Graph. Cumulative distribution function plot for all GRS composite tests.

Table 24. Predicted and measured vertical capacity for all GRS composite tests.

Test

Measured Predicted Bias, 

Measured

Predicted

Standard 

Normal 

Variable, 

z

Reference

No.

qult,emp 

(psf)

qult,an,c

(psf)

GSGC 2 70,957 51,841 1.37 1.24 11

GSGC 3 42,574 40,207 1.06 0.46

GSGC 4 37,252 26,795 1.39 1.40

GSGC 5 42,449 46,485 0.91 -0.10

Elton 1 4,805 5,324 0.90 -0.17 8

Elton 2 2,695 3,178 0.85 -0.31

Elton 3 6,392 6,909 0.93 0.03

Elton 4 6,100 7,226 0.84 -0.38

Elton 5 8,398 8,495 0.99 0.31

Elton 6 8,293 8,812 0.94 0.10

Elton 7 9,589 10,397 0.92 -0.03

NCHRP 1 8,356 5,685 1.54 1.93 33

NCHRP 2 17,757 18,287 0.97 0.24

Defiance 1 13,370 11,797 1.13 0.70 1

Defiance 2 25,068 23,030 1.09 0.61

Bathurst 1,670 1,283 1.30 1.10 34

λ

y = 3.33x - 3.16

R2 = 0.98

-3

-2

-1

0

1

2

3

0 0.5 1 1.5 2 Standard Normal Veriable, z

Bias, λ
84

MP A 4,696 6,371 0.74 -0.70 4

MP B 3,548 14,044 0.25 -1.93

MP C 9,600 13,493 0.71 -0.79

DC-1 23,310 30,439 0.77 -0.32 N/A

DC-2 22,709 25,952 0.88 0.16

DC-5 21,539 30,439 0.71 -0.49

TF-1 20,487 16,447 1.25 1.15

TF-2 25,260 23,403 1.08 0.67

TF-3 17,491 22,741 0.77 -0.16

TF-5 25,920 44,786 0.58 -1.56

TF-6 43,763 45,448 0.96 0.38

TF-7 26,546 44,786 0.59 -1.19

TF-8 29,134 44,786 0.65 -0.93

TF-9 22,310 15,369 1.45 1.56

TF-10 10,330 14,707 0.70 -0.54

TF-11 23,249 28,348 0.82 0.07

TF-12 29,030 29,001 1.00 0.54

TF-13 12,960 18,764 0.69 -0.72

TF-14 23,562 19,426 1.21 0.93

Mean Bias 0.95

Standard Deviation 0.30

Coefficient of Variation (percent) 0.32

N/A = not applicable, no reference number is associated with a test since results are 

being reported for the first time in this report.

7.2.2 Empirical

The strength design limit for capacity is defined as the stress at which the GRS composite fails 

(meaning it cannot sustain any additional load) or strains 5 percent vertically, whichever occurs 

first. To relate this to the ultimate capacity and to non-dimensionalize the results, the stress-strain 

curves were normalized by the ultimate capacity of the particular GRS composite for all of the 

PTs taken to failure (figure 81). The 5-percent vertical strain limit can significantly reduce the 

allowable stress placed on a GRS composite; figure 82 shows that at 5-percent strain, the applied 

pressure is between 50 and 85 percent of the measured ultimate capacity (table 8).

Along with the additional factor of safety of 3.5, using 5-percent vertical strain as a design 

strength limit for the empirical internal stability design is a conservative measure; however, it 

accounts for the resources available at most laboratory institutions so that PTs can be terminated 

at a reasonable, yet achievable, strain limit. The more important design limit for GRS walls and 

abutments is the service limit.
85

Figure 81. Graph. Normalized applied stress versus strain for all PT. 
86

Figure 82. Graph. Normalized load-deformation behavior for the DC and TF PTs up to 

5-percent vertical strain.
87

7.3 SERVICE LIMIT

FHWA guidance currently recommends a service limit of 4,000 psf applied vertical stress (dead 

plus live load, with load factors = 1.0) with a target of 0.5-percent vertical strain for GRS 

abutments, measured by using only dead load, unless the engineer decides to permit additional 

load and/or deformation.(1)For a 20-ft-tall wall, the initial compression of the GRS composite 

would therefore be limited to 1.2 inches. Note that the majority of compression within a GRS 

abutment occurs shortly after placement of dead load, with little post-construction settlement, if 

constructed as outlined by FHWA.(5)

The stress-strain curve generated by the PT can immediately be used to estimate deformation. By 

knowing the applied dead load, the resulting strain can be found from the PT results. 

Alternatively, using a corresponding PT for the particular GRS composite would yield the 

amount of dead load (unfactored) that can be placed on the abutment to limit the deformation to 

the tolerable amount. For the range of tests conducted in this study (compacted only), the 

allowable applied stress is between 1,324 and 6,016 psf (table 8 and figure 84).

Using the normalized load test curves (figure 85), the service stress to meet the 0.5-percent

vertical strain limit is 6 to 23 percent of the ultimate stress, with an average of 13 percent for all 

tests. For only the tests with a CMU facing, as would be typical in the field, the range is slightly 

narrower, between 8 and 19 percent with an average of 12 percent; for the tests without a facing 

element, the average is 15 percent.

When looking strictly at the PTs that meet the current FHWA specifications for a GRS abutment, 

the allowable applied stress at low strain levels (i.e., at in-service conditions) ranges between 

about 8 percent and 13 percent of the ultimate stress (figure 86), with an average of 10 percent.

(1)

By estimating an ultimate stress using the GRS capacity equation (figure 8), ignoring the effects 

of cohesion and confining stress as indicated by Adams et al. (2011a) for design of GRS 

abutments, and assuming 10 percent of the ultimate capacity, the maximum service stress 

(unfactored dead load only) to limit short-term vertical strain to 0.5 percent can be predicted

(table 25).(1)
88

Figure 83. Graph. Cumulative distribution function for proposed service limit pressure.

y = 1.71x - 1.97

R2 = 0.95

-2

-1.5

-1

-0.5

0

0.5

1

1.5

2

0 0.5 1 1.5 2 2.5

Standard Normal Veriable, z

Bias, λ
89

Table 25. Estimation of allowable dead load to limit vertical strain to 0.5 percent using the 

GRS capacity equation.

Test

Measured Predicted Bias, 

Measured

Predicted

Standard 

Normal 

qult,emp (psf) (psf) Variable, z

, predicted =

10% of qult,an (psf)

DC-1 23,310 3,065 2,974 1.03 -0.34

DC-2 22,709 2,171 2,551 0.85 -0.48

DC-3 N/A 1,324 4,021 0.33 -1.62

DC-4 N/A 2,212 1,655 1.34 0.48

TF-1 20,471 2,068 1,570 1.32 0.20

TF-2 25,260 4,759 2,204 2.16 1.62

TF-3 17,491 3,417 2,204 1.55 1.00

TF-4 N/A 3,705 4,409 0.84 -0.80

TF-5 25,920 6,016 4,409 1.36 0.63

TF-6 43,763 3,704 4,409 0.84 -0.63

TF-7 26,546 2,224 4,409 0.50 -1.00

TF-8 29,134 1,753 4,409 0.40 -1.25

TF-9 22,310 2,955 1,401 2.11 1.25

TF-10 10,330 1,586 1,401 1.13 -0.07

TF-11 23,249 2,839 2,765 1.03 -0.20

TF-12 29,030 4,028 2,765 1.46 0.80

TF-13 12,960 2,398 1,807 1.33 0.34

TF-14 23,562 2,037 1,807 1.13 0.07

Mean Bias 1.15

Standard Deviation 0.58

Coefficient of Variation (percent) 0.51

N/A = Not applicable, did not fail the composite.

The results indicate that by using 10 percent of the predicted design capacity, the mean bias 

between the measured results from this series of PTs and the predicted allowable stress at 

0.5-percent vertical strain is 1.15 with a COV of 0.51 (table 25). This is on the conservative side 

and offers another tool to estimate deformation in lieu of conducting a performance test on a 

particular GRS composite.

λ

q@ε=0.5% q@ε=0.5%
90

Figure 84. Graph. Load-deformation behavior for the Turner Fairbank PTs at low strain 

levels.
91

Figure 85. Graph. Normalized load-deformation behavior for the DC and TF PTs up to 

0.5-percent vertical strain.
92

Figure 86. Graph. PTs strictly meeting FHWA GRS abutment design specifications.
93

7.4 LRFD CALIBRATION FOR STRENGTH LIMIT

In the United States, bridges that receive federal funding for construction must be designed in 

accordance with the AASHTO LRFD Bridge Design Specifications. AASHTO first adopted a 

load factor design (LFD) in 1977 for certain portions of the bridge superstructure. In the mid1990s, the AASHTO Specifications adopted LRFD in geotechnical engineering design based on 

the work of Barker et al. (1991).(35)Since then, LRFD research has extended to more and more 

aspects of geotechnical structures.

LRFD in geotechnical design of bridge and building foundations has the following benefits:

1. Efficiency in design effort can be achieved when using the same factored loads for both 

superstructure and sub-structure design.

2. More consistent margins of safety can be realized in both superstructure and sub-structure 

design.

3. More economical use of materials can be accomplished when a more rational basis for setting 

the margin of safety is provided.

A methodology for calibrating the load bearing capacity (strength limit state) of bridge 

foundations on a GRS abutment is described and presented. Calibration for serviceability limit 

state is just as, if not more important, since most shallow foundation designs are governed more 

by serviceability rather than the ultimate limit state; however, the scope of this report is limited 

to only the strength limit state.(10)

7.4.1 Background

Code calibration is the process of determining values of load and resistance factors. While load 

factors have already been well established in the AASHTO code, resistance factors for newer 

types of geotechnical structures must be obtained by calibration. Resistance factors can be 

calibrated using the following approaches:

1. Judgment.

2. Fitting with Allowable Stress Design (ASD).

3. Reliability theory.

Calibration by judgment was the first approach at arriving code parameters, but it suffers from 

non-uniform margins of safety. Calibration by fitting with ASD essentially results in the LRFD 

code mimicking the ASD code. It does not necessarily result in more uniform margins of safety 

or economy. Calibration by fitting was performed for the internal stability design of GRS 

abutments.

(1) The approach adopted herein to calibrate the load bearing capacity of GRS 

composites is based on reliability theory, which offers a rational process for optimizing the value 

of resistance factor. The results are then compared with those from calibrating by fitting with 

ASD.

In general, there are three approaches to estimating resistance factors using reliability analysis: 

(1) first order second moment (FOSM); (2) advanced first order second moment (AFOSM) and 

(3) Monte Carlo simulation (MCS). FOSM is the simplest and results in resistance factors that 
94

are marginally lower (more conservative) than the AFOSM method, which requires 

programming of an optimization routine. MCS is only accurate if the load and resistance 

distributions are correctly modeled. According to Paikowsky (2010), if the statistical parameters 

are not well-defined, the resulting resistance factor using MCS will not be correct; moreover, if 

the load and resistance distribution is assumed to be lognormal, MCS will theoretically produce 

the same results as the FOSM method.(36) Thus, the simpler FOSM approach is adopted in this 

report.

In the reliability-based approach, the load (Q) and resistance (R) are considered random 

variables, which can be described by their own probability density functions. One form of 

margin of safety can be expressed by the difference R – Q where failure occurs when R – Q < 0. 

In the FOSM approach, R and Q are assumed to have a lognormal distribution; hence a limit 

state function can be written (figure 87) which is also lognormal. Failure occurs when R/Q is less 

than 1 or when g(R,Q) is less than 0. The probability density function of figure 87 is shown in 

figure 88.

Figure 87. Equation. Limit state function for FOSM approach.

Figure 88. Graph. Reliability index for lognormal R and Q.

The reliability index ( ) is the number of standard deviations between the mean safety margin

(ḡ) and failure, where g(R,Q) = 0. According to Paikowsky (2010), many shallow foundations 

supporting bridge elements have been designed with a target reliability index ( ) between 3.0 

and 3.5, which corresponds to a probability of failure of 0.135 percent and 0.023 percent, 

respectively.(36) Note that these target reliability indices are for foundations on natural soil 

whereas GRS is an engineered fill.

7.4.2 Reliability Analysis: FOSM

In LRFD, the factored resistance must be greater than or equal to the factored loads (figure 89).

β

βT

g(R,Q) = In(R/Q) 
95

Figure 89. Equation. LRFD format.

Where = resistance factor, = load factor for load component i, R = resistance and Qi = load 

component i (e.g., dead load or live load). Many load combinations exist in the AASHTO code; 

however, resistance factors herein are calibrated using the Strength I load combination only. 

With only dead and live loads considered, it can be shown that the resistance factor based on 

FOSM can be calculated according to figure 90.

Figure 90. Equation. Resistance factor using FOSM.

Where , , and are bias factors for resistance, dead load and live load, respectively,

and are the dead (1.25 for AASHTO assuming Strength I load combination and assuming 

dead load of structural components and non-structural attachments only; dead load due to the 

wearing surface is not considered as it usually represents only a small fraction of the total dead 

weight) and live (1.75 for AASHTO assuming Strength I load combination) load factors, 

respectively, and VR and VQ are the coefficients of variation (COV) of the resistance and loads, 

respectively. The bias factor, , is the ratio of the measured to the nominal or calculated value; if 

the variable tends to be over-predicted, then the bias factor will be less than 1. Note that the COV

for a variable that is related to several statistically independent variables is merely the square 

root of the sum of the squares of the various coefficients of variation. For example, figure 91

shows that the COV for factored load (VQ) is equal to the sum of the squares of the COV for 

dead load (VD) and live load (VL).

Figure 91. Equation. Coefficient of variation for factored load.

It can be seen that in the FOSM method to estimate the resistance factor (figure 90), the random 

variables are characterized by their first two moments, namely the mean and the standard 

deviation (which is equal to the product of the coefficient of variation and the mean). Therefore, 

it is essential to compile the statistics of the load and resistances prior to calibration.

ΦR ≥ γiQi

n

i=1

Φ i

Φ =

λr γD 

QD

QD + QL

 / 1 −

QD

QD + QL

 + γL 

 γD 

QD

QD + QL

 / 1 −

QD

QD + QL

 + γL 

1 + VR

2

1 + VQ

2

e

 βT ln 1+VR

2

 1+VQ

2

 

i

R D L

D L

λ

VQ = VD

2 + VL2

2
96

Load Statistics 

In the final report to the National Cooperative Highway Research Program (NCHRP) Project 

No. 20-7/186, Kulicki et al. (2007) provided bias factors and COVs for dead and live loads 

(table 26).(37) Thus, from (figure 91) and table 26, the COV for factored load is equal to 

15.6 percent. 

Table 26. Statistics for dead and live loads. 

Dead Load1 Live Load2

Bias Factor = 1.05 = 1.14

COV VD = 10% VL = 12%

1Assumes cast-in-place concrete elements only. 

2Assumes HL-93 loading for shear and an average daily truck traffic 

(ADTT) of 1,000 for two loaded lanes.(10,37)

Resistance Statistics 

The variables in the soil-geosynthetic capacity equation (figure 8) and the required reinforcement 

strength equation (figure 10) include the maximum aggregate size (dmax), the reinforcement 

strength (Tf) and the passive earth pressure coefficient (Kpr); it is reasonable to assume that the 

reinforcement spacing (Sv) is not a random variable, therefore the associated fabrication factor is 

assumed to be unity. Therefore, the COV of the resistance (VR) can be expressed according to 

figure 92. 

Figure 92. Equation. Coefficient of variation for resistance. 

Where Vdmax, VTf, VKp and VM are the COVs of dmax, Tf, Kp and the capacity model, respectively. 

Note that the ultimate reinforcement strength is usually measured in the laboratory under 

unconfined conditions; however, it should ideally be measured under confinement to replicate 

actual in-service conditions. Benjamin et al. (2008) performed numerous measurements of 

reinforcement strength for woven and non-woven geotextiles in the ground and found the COV 

for reinforcement strength (VTf) of woven geotextiles was 2.3 and 4.25 percent for the crossmachine and machine directions, respectively.(38) Since woven geotextiles were used in the PTs 

and are the most commonly used geosynthetic in GRS abutments and IBSs, a VTf value of 

4 percent is utilized in this study. For the COV for the passive earth pressure coefficient (VKp), 

NCHRP Report 651 recommended a value of 15 percent when measuring friction angles with lab 

tests.(36) The COV for maximum aggregate size (Vdmax) has not yet been investigated, but it is 

expected to be small; for one reason, the quarries will set their crushers to a given maximum 

grain size with a high degree of certainty. A value of Vdmax = 5 percent is therefore assumed in 

this calibration. Additionally, the mean bias factor for the capacity prediction model, using only 

the data produced in this study, is 0.88 with a COV of about 35 percent (table 23). The resulting 

COV for resistance (VR) is estimated at 39 percent. 

D L

VR = Vdmax

2 + VTf2 + VKp2 + VM

2 2
97

7.4.3 Resistance Factor

The database of results from this study consists of sixteen cases taken to failure of the GRS 

composite, four with open-graded backfill (DC-1, DC-2, DC-5, and TF-1) and twelve with wellgraded backfill material (TF-2, TF-3, and TF-5 to TF-14). Because the number of cases with 

open-graded backfill is limited, the present strategy is to arrive at one value for the resistance 

factor that is applicable to all conditions. With additional testing in the future, separate resistance 

factors can be calibrated for each soil type, if necessary.

The resistance factor for GRS capacity, calculated according to figure 90, is presented as a 

function of the ratio of dead load to total load, QD/(QD + QL), and the target reliability index ( ) 

in figure 93. Note that a QD/(QD + QL) of 0.5 corresponds approximately to a bridge span of 

about 60 ft, whereas a QD/(QD + QL) of 0.2 corresponds to a bridge span of approximately 

250 ft.

(39) The bridge span for a GRS-IBS is currently limited to 140 ft (Adams et al. 2011);

therefore the resistance factors of interest correspond to QD/(QD + QL) greater than about 0.32. 

Over this range, the resistance factor varies from 0.26 to 0.28 for = 3.5, from 0.32 to 0.34 for 

= 3.0, and from 0.40 to 0.41 for = 2.5 using FOSM.

Figure 93. Graph. Resistance factor for footings on GRS composites for different dead to 

dead plus live load ratios and target reliability indices based on PT series.

Similarly, a reliability analysis was performed using the results of all prior GRS composite 

testing with a mean bias of 0.95 and a COV for the model of 32 percent (table 24). This slightly 

increases the resistance factor for the same target reliability indices (figure 94); for the range of 

QD/(QD + QL) greater than 0.32, the resistance factor varies from 0.31 to 0.33 for = 3.5, from 

0.38 to 0.39 for = 3.0, and from 0.46-0.48 for = 2.5 using FOSM.

βT

βT

βT βT

βT

βT βT
98

The resistance factors found through the reliability analysis using the standard reliability indices 

of 3.0 and 3.5 for bridge foundations are lower than the resistance factor found through 

calibration by fitting to ASD. For a factor of safety (FS) of 3.5, a resistance factor of 0.45 was 

estimated for the soil-geosynthetic capacity equation.(1) Through back-calculation, this equates to 

a reliability index ( ) of about 2.5.

In NCHRP Report 507, Paikowsky et al. (2004) suggested that for redundant foundation systems 

(such as pile groups), a reliability index of 2.3 was appropriate.(40) With the frequency of the 

reinforcement spacing, GRS composites can be considered a redundant foundation system as 

well, with no catastrophic collapse observed at failure for any of the PTs. In addition, Bathurst et 

al. (2008) suggest that a reliability index of 2.3 is appropriate for the internal stability of 

reinforced soil walls.(41) This leads to the conclusion that the currently used resistance factor of 

0.45, having a reliability index of 2.5, is reasonable for design.

Figure 94. Graph. Resistance factor for footings on GRS composites for different dead to 

dead plus live load ratios and target reliability indices based on all testing to date.

β
99

8. CONCLUSIONS

The results of the PT research performed on GRS composites leads to several important 

conclusions. General conclusions include the following:

1. The performance test can be used to model the load-deformation behavior of a particular 

GRS composite and is repeatable. Based on equivalency comparisons, the ratio of the 

measured vertical capacity in a PT to that of the same GRS in plane strain is near unity, and 

the stiffness of what a plane strain GRS composite might experience (such as an abutment) is 

up to 3 times higher (on the conservative side) compared to what is measured in a GRS PT

(section 5).

2. A two-post frame with single hydraulic jacks produces more even load distribution than a 

two bolted channel beams in combination with four hydraulic jacks.

The load-deformation relationship of GRS composites depends on several parameters including 

preloading, aggregate angularity, compaction level, presence of bearing bed reinforcement, and 

facing confinement.

 Figure 32 indicates that preloading of the GRS composite (TF-6 with Sv = 75⁄8 inches, 

Tf = 4,800 lb/ft, 21A material, CMU facing) resulted in strain hardening to produce a stiffer 

response during two unload-reload cycles. The reload response is similar during both cycles, 

with a ratio of stress to strain equal to approximately 3,000 ksf, compared to the ratio for the 

initial, virgin loading equal to approximately 750 ksf, an increase by a factor of four. Similar 

results were found for an open-graded material as well (TF-1 with Sv = 75⁄8 inches, 

Tf = 2,400 lb/ft, AASHTO No. 8 material, CMU facing), as shown in figure 30. The 

effect of preloading was not explored for other reinforcement spacing in this study.

 Angularity impacts the strength (e.g., friction angle) of the backfill used and therefore the 

GRS composite. The GRS composite using a rounded pea gravel aggregate (DC-2) had both 

a lower ultimate strength (qult,emp) and service limit capacity ( ) than that using the 

more angular aggregate (DC-1) meeting the same gradation specifications for an AASHTO 

No. 8 material (table 10).

 The results indicate a similar ultimate vertical capacity (qult,emp) between an uncompacted

(DC-5) and compacted (DC-1) GRS composite, but a much softer response with no 

compactive effort. For the particular GRS composite tested (table 11 and figure 58), at an 

in-service dead load of 4,000 psf, an uncompacted abutment will experience an initial strain 

about three times more than a compacted abutment; however, near failure, at about 

20,000 psf, an uncompacted abutment will experience about 1.7 times the strain as a 

compacted abutment.

 The results indicate that the bearing bed provides some added vertical capacity; however, 

vertical deformation is not improved at low strain levels (figure 60). The modulus for 

primary compression is similar whether the bearing bed is present or not (table 12). At both 

low and high applied normal stresses, representing bridge loads (figure 61) and near failure 

loads (figure 62), respectively, the bearing bed reinforcement serves to limit lateral 

deformation in the zone of its placement.

q@ε=0.5%
100

 The frictionally connected CMU facing has an impact on the performance of GRS; it 

provides confinement, leading to a stiffer response and an increased capacity compared to a 

GRS composite with no facing element. From table 20, the facing more than doubles the 

initial stress-strain ratio as compared to the PTs without any facing. In terms of capacity, the 

facing plays the biggest and smallest role for the largest spaced (Sv = 151⁄4 inches) and the 

closest spaced (Sv = 313/16 inches) system tested, respectively. For the 5 pairs of tests 

conducted at TF, including the CMU facing produced an improved ultimate capacity between 

1.2 and 2.2 times greater than the GRS composite without any facing (table 21). A similar 

trend to the strain at the current 4,000-psf service limit is also found. The design assumption 

to not include the effect of confinement from the face in determining the capacity and 

required reinforcement strength is therefore conservative.(1)

 In addition, at the current service stress limit (applied stress, q, of 4,000 psf), the ratio of 

service vertical strain with no facing to service vertical strain with a CMU facing ranges from 

1.2 to 2.2 (table 21). The largest impact was for the 155⁄8-inch spaced GRS composite (TF-9 

and TF-10). At ultimate failure, the ratio is considerably closer, ranging from 0.8 to 1.2.

 For the same Tf/Sv ratio of 3,800 lb/ft2, there is a linear relationship between the 

reinforcement spacing (Sv) and the ratio of capacity with CMU facing (qult,emp CMU) to the 

capacity without facing (qult,emp no CMU) (figure 75). Similarly, there is a linear relationship 

with reinforcement strength (figure 76).

Both open-graded and well-graded aggregates can be used as the reinforced backfill in GRS 

composites; each has their advantages and disadvantages. The PT results further distinguished 

the effect aggregate selection can have on the behavior of GRS.

 The well-graded material is considerably stiffer than the open-graded material. At an applied 

pressure of 4,000 psf, the vertical strain is about 1.1 percent for TF-1 (open-graded) and 

0.4 percent for TF-2 (well-graded). The modulus of the composite tested in TF-1 is 320 ksf 

(table 13) compared to 710 ksf for the composite tested in TF-2; TF-1 (open-graded) is 

55 percent less stiff than TF-2 (well-graded). In terms of bearing capacity, TF-1 was 

20 percent less strong than TF-2. The results indicate that the gradation, and perhaps 

cohesion (at the same friction angle), impacts the modulus more so than the strength, giving 

indication that well-graded fills have an advantage with respect to serviceability.

 As discussed in section 6.1, isolating the effect of cohesion and other soil parameters on the 

performance of GRS is difficult using PTs, although based on the soil-geosynthetic capacity 

equation (figure 8), it will serve to improve capacity, although its contribution should not be 

considered in design.

 The response indicates that the higher reinforcement strength (4,800 lb/ft) produces a stiffer 

and stronger response than the lower reinforcement strength (2,400 lb/ft) for open-graded

backfill (table 15); the same is true for the composites with no facing element. By doubling 

the reinforcement strength, the results indicate an increase of measured capacity by a factor 

of 1.14 for capacity and 1.34 for the initial stress-strain ratio.

 For the well-graded aggregate, increasing the reinforcement strength increases the capacity 

by a factor of 1.5 and 1.7 for no facing and CMU facing (table 16), respectively, but does not 

significantly impact the stiffness.
101

The relationship between reinforcement strength and spacing was investigated through a series of 

PTs with the same Tf/Sv ratio.

 The response for experiments conducted with a Tf/Sv ratio of 3,800 lb/ft2showed that as 

reinforcement spacing increased, the vertical capacity decreased for the same Tf/Sv ratio, 

whether a CMU facing was absent (figure 66) or present (figure 67). This suggests that the 

relationship between reinforcement strength and spacing is not proportional to capacity as

outlined in current MSE design (AASHTO 2012); a GRS abutment with a given Tf and Sv

will not have the same strength as a GRS abutment with twice the strength (2Tf) and 

reinforcement spacing (2Sv).

 At the same Tf/Sv ratio of 3,800 lb/ft2, increasing the reinforcement spacing by a factor of 

two from 75⁄8 to 151⁄4 inches and the reinforcement strength by a factor of two from

2,400 to 4,800 lb/ft, resulted in a reduction of the capacity by a factor of 0.9 and 0.6 for CMU 

facing and no facing, respectively. The relationship is therefore not directly proportional, as 

indicated for MSE design (AASHTO 2012).

The primary purpose of PTs is to provide a designer with the unique stress-strain properties of a 

particular GRS composite for use in design. There is an empirical method and an analytical 

method currently available for GRS by FHWA.

(1)

 Note that only 3 GRS composites tested in this study meet both the design limits and the 

material specifications (e.g., Tf ≥ 4,800 lb/ft, dmax ≥ 1⁄2 inches) for GRS abutments; however, 

based on the results of this study, perhaps the criteria can be amended to reduce the 

reinforcement strength limitation, provided the reinforcement strength meets internal stability 

design requirements.(1)

 Based on the results of the PTs presented in this report, the average bias for the capacity 

equation is 0.88 with a COV of about 35 percent (table 23). Including the results from other 

sources in the literature, the average bias is 0.95 with a COV of about 32 percent. These 

numbers indicate relatively good agreement between figure 8 and the measured ultimate 

capacity (table 24).

 The 5-percent vertical strain limit can significantly reduce the allowable stress placed on a 

GRS composite; figure 82 shows that at 5-percent strain, the applied pressure is between 50 

and 85 percent of the measured ultimate capacity (table 8).

 The results indicate that by using 10 percent of the predicted design capacity (figure 8), the 

0.5-percent vertical strain as required by FHWA for the service limit of GRS abutments can 

largely be satisfied. The mean bias of the ratio between the measured results at 0.5-percent

vertical strain from this series of PTs and the predicted allowable stress at 10 percent of the 

design capacity is 1.15 with a COV of 0.51 (table 25). This is on the conservative side yet

offers another tool to estimate deformation in lieu of conducting a performance test on a 

particular GRS composite.

Finally, a reliability analysis for the soil-geosynthetic capacity equation (figure 8) was performed 

on the 16 PTs taken to failure in this study, along with previous results from additional GRS 

testing found from the literature. 
102

 A target reliability index of 2.5 is reasonable for the strength limit models for GRS 

composites designed according to Adams et al. (2011). A lower target reliability index is 

warranted for closely-spaced GRS composites because of the redundancy in the 

reinforcement and because no catastrophic collapse was observed in any of the performance 

tests at failure. In addition, Bathurst et al. (2008) suggest that a reliability index of 2.3 is 

appropriate for the internal stability of reinforced soil walls.(1,41)

 Performing a reliability analysis using the FOSM approach produces a resistance factor 

around 0.45, similar to that found through calibration by fitting to ASD methods.
103

APPENDIX A. SOIL TESTING DATA

SIEVE ANALYSIS DATA

Table 27. AASHTO No. 8 sieve analysis (DC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (g)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

0.75 0.75 0 1,393.3 0 0.00 100.00

0.50 0.5 0 1,087.9 0 0.00 100.00

0.375 0.375 1,314.4 1,102.7 211.7 3.01 96.99

4 0.187 6,469.6 1,515 4,954.6 70.49 26.50

8 0.0937 2,666.2 1,129 1,537.2 21.87 4.63

16 0.0469 1,105.3 910.2 195.1 2.78 1.85

50 0.0117 807.6 741.5 66.1 0.94 0.91

100 0.0059 820.1 809.4 10.7 0.15 0.76

200 0.0029 924.1 916.4 7.7 0.11 0.65

Pan — 715.4 669.8 45.6 0.65 0.00

Table 28. AASHTO No. 8 pea gravel sieve analysis (DC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (kg)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

0.75 0.75 0 0 0 0.00 100.00

0.50 0.5 1.115 1.085 0.03 0.43 99.57

0.375 0.375 1.38 1.1 0.28 3.99 95.58

4 0.187 6.7 1.015 5.685 80.98 14.60

8 0.0937 1.665 1.13 0.535 7.62 6.98

16 0.0469 1.095 0.91 0.185 2.64 4.34

50 0.0117 0.86 0.74 0.12 1.71 2.64

Pan — 0.785 0.6 0.185 2.64 0.00
104

Table 29. AASHTO No. 57 Sieve analysis (DC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (g)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

1.5 1.5 1,476.1 1,476.1 0 0.00 100.00

1 1 1,127.1 1,127.1 0 0.00 100.00

0.75 0.75 2,311 1,393.4 917.6 12.09 87.91

0.50 0.5 5,052.1 1,087.8 3,964.3 52.22 35.69

0.375 0.375 2,739.5 1,102.7 1,636.8 21.56 14.13

4 0.187 2,308.8 1,514.8 794 10.46 3.67

8 0.0937 1,224.5 1,128.9 95.6 1.26 2.41

16 0.0469 981.1 909.8 71.3 0.94 1.47

200 0.0029 974.5 916.7 57.8 0.76 0.71

Pan — 723.3 669.7 53.6 0.71 0.00

Table 30. AASHTO No. 9 Sieve analysis (DC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (g)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

0.75 0.75 0 1,393.21 0 0.00 100.00

0.50 0.5 0 1,087.95 0 0.00 100.00

0.375 0.375 0 1,102.79 0 0.00 100.00

4 0.187 1,785.33 1,514.98 270.35 5.78 94.22

8 0.0937 4,236.2 1,129.04 3,107.16 66.47 27.75

16 0.0469 1,791.1 909.95 881.15 18.85 8.90

50 0.0117 986.6 741.69 244.91 5.24 3.66

100 0.0059 830 809.4 20.6 0.44 3.22

200 0.0029 934.9 916.46 18.44 0.39 2.82

Pan — 732.6 600.56 132.04 2.82 0.00
105

Table 31. AASHTO No. 8 Sieve analysis (TFHRC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (g)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

1 1 1,127.1 1,127.1 0 0.00 100.00

0.50 0.5 1,116.5 1,087.7 28.8 0.31 99.69

0.375 0.375 3,832 1,102.6 2,729.4 29.83 69.86

4 0.187 6,694.3 1,013.8 5,680.5 62.08 7.78

8 0.0937 1,688.5 1,128.7 559.8 6.12 1.66

10 0.079 1,398.7 1,374 24.7 0.27 1.39

16 0.0469 935 909.7 25.3 0.28 1.11

40 0.0165 810 793.1 16.9 0.18 0.93

50 0.0117 746.2 741.8 4.4 0.05 0.88

Pan — 750.4 669.8 80.6 0.88 0.00

Table 32. AASHTO A-1-a (VDOT 21A) sieve analysis (TFHRC tests).

Sieve 

No.

Sieve Size 

(inch)

Mass of Soil + 

Pan (g)

Mass of Pan 

(kg)

Mass of 

soil (kg)

Percent

Retained

Percent

Passing

1 1 1,127.1 1,127.1 0 0.00 100.00

0.50 0.5 1,919.2 1,087.6 831.6 17.59 82.41

0.375 0.375 1,624.9 1,102.7 522.2 11.05 71.36

4 0.187 2,093.7 1,013.8 1,079.9 22.84 48.52

8 0.0937 1,757.5 1,129.6 627.9 13.28 35.24

10 0.079 1,489.5 1,374.7 114.8 2.43 32.81

16 0.0469 1,260.5 910.1 350.4 7.41 25.40

40 0.0117 1,206.5 793.1 413.4 8.74 16.66

200 0.0029 1,397.8 916.5 481.3 10.18 6.47

Pan — 975.7 669.6 306.1 6.47 0.00
106

LSDS TESTING DATA

Table 33. Summary of AASHTO No. 8 LSDS results (DC tests).

Soil 

Type

Normal 

Stress (psi)

Max Shear 

(psi)

Peak Friction 

Angle (°)

Peak Dilation 

Angle (°)

Gap 

(inch)

8 5.051 9.5114 62.03 12.96 0.343

8 10.03 19.0619 62.25 11.65 0.343

8 20.03 30.533 56.73 6.35 0.343

8 30.02 45.6537 56.67 5.63 0.343

Figure 95. Graph. AASHTO No. 8 LSDS test results (DC tests).

0

5

10

15

20

25

30

35

40

45

50

0 0.5 1 1.5 2 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
107

Figure 96. AASHTO No. 8 LSDS deformation test results (DC tests).

Table 34. Summary of AASHTO No. 8 pea gravel LSDS results (DC tests).

Soil 

Type

Normal 

Stress (psi)

Max Shear 

(psi)

Peak Friction 

Angle (°)

Peak Dilation 

Angle (°)

Gap 

(inch)

8P 5.022 7.5449 56.35 13.19 0.35

8P 10.05 15.00489 56.19 11.79 0.35

8P 20.01 21.238 46.71 5.96 0.35

8P 30.01 35.1467 49.51 6.04 0.35

-0.25

-0.2

-0.15

-0.1

-0.05

0

0.05

0 0.5 1 1.5 2 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
108

Figure 97. Graph. AASHTO No. 8 pea gravel LSDS test results (DC tests).

Figure 98. Graph. AASHTO No. 8 pea gravel LSDS deformation test results (DC tests).

-0.25

-0.2

-0.15

-0.1

-0.05

0

0.05

0 1 2 3 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi

0

5

10

15

20

25

30

35

40

0 1 2 3 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
109

Table 35. Summary of AASHTO No. 57 LSDS results (DC tests).

Soil 

Type

Normal 

Stress (psi)

Max Shear 

(psi)

Peak Friction 

Angle (°)

Peak Dilation 

Angle (°)

Gap 

(inch)

57 5.013 17.547 74.06 15.36 0.74

57 10.01 21.187 64.71 13.15 0.74

57 20.04 40.212 63.51 9.93 0.74

57 30.01 47.189 57.55 6.28 0.74

Figure 99. Graph. AASHTO No. 57 LSDS test results (DC tests).

0

5

10

15

20

25

30

35

40

45

50

0 1 2 3 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
110

Figure 100. Graph. AASHTO No. 57 LSDS deformation test results (DC tests).

Table 36. Summary of AASHTO No. 9 LSDS results (DC tests).

Soil 

Type

Normal 

Stress (psi)

Max Shear 

(psi)

Peak Friction 

Angle (°)

Peak Dilation 

Angle (°)

Gap 

(inch)

9 5.053 9.7334 62.56 13.64 0.175

9 10.04 17.698 60.43 10.38 0.175

9 20.01 31.97 57.96 5.82 0.175

9 30.02 42.8381 54.98 4.73 0.175

-0.35

-0.3

-0.25

-0.2

-0.15

-0.1

-0.05

0

0.05

0 1 2 3 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
111

Figure 101. Graph. AASHTO No. 9 LSDS test results (DC tests).

Figure 102. Graph. AASHTO No. 9 LSDS deformation test results (DC tests).

0

5

10

15

20

25

30

35

40

45

0 1 2 3 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi

-0.2

-0.15

-0.1

-0.05

0

0.05

0 1 2 3 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
112

Table 37. Summary of AASHTO No. 8 LSDS results (TFHRC tests).

Soil 

Type

Normal 

Stress (psi)

Max Shear 

(psi)

Peak Friction 

Angle (°)

Peak Dilation 

Angle (°)

Gap 

(inch)

8 5.04 15.220948 71.68 15.66 0.443

8 10.06 26.73174 69.38 11.01 0.443

8 20.03 39.8865 63.34 8.07 0.443

8 30.02 52.54942 60.26 5.10 0.443

Figure 103. Graph. AASHTO No. 8 LSDS test results (TFHRC tests).

0

10

20

30

40

50

60

0 1 2 3 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
113

Figure 104. Graph. AASHTO No. 8 LSDS deformation test results (DC tests).

Table 38. Summary of AASHTO A-1-a (VDOT 21A) LSDS results (TFHRC tests).

Soil

Type

Test

Condition

Normal

Stress

(psi)

Max

Shear

(psi)

Peak

Friction

Angle (°)

Peak

Dilation

Angle (°)

Gap (inch)

21A Dry 5.046 17.305 73.74 26.08 0.58

21A Dry 10.05 23.6762 67.00 20.48 0.58

21A Dry 20.02 40.35021 63.61 12.58 0.58

21A Dry 30.03 50.135 59.08 11.05 0.58

21A Saturated 4.99 5.46532 47.60 0.00 0.58

21A Saturated 9.991 14.2569 54.98 0.05 0.58

21A Saturated 19.98 25.034 51.41 0.00 0.58

21A Saturated 29.99 36.2423 50.39 0.00 0.58

-0.4

-0.35

-0.3

-0.25

-0.2

-0.15

-0.1

-0.05

0

0.05

0 1 2 3 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
114

Figure 105. Graph. AASHTO A-1-a (VDOT 21A) LSDS test results (TFHRC tests).

Figure 106. Graph. AASHTO A-1-a (VDOT 21A) LSDS deformation test results (DC tests).

0

10

20

30

40

50

60

0 0.5 1 1.5 2 2.5 3 Shear Stress (psi)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi

-0.5

-0.4

-0.3

-0.2

-0.1

0

0.1

0 1 2 3 Vertical Displacement (inch)

Horizontal Displacement (inch)

5 psi

10 psi

20 psi

30 psi
115

APPENDIX B. NUCLEAR DENSITY TESTING FOR TFHRC PTS

Table 39. TF-2 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 backscatter 107.8 95.6 1.2

backscatter 109.2 96.8 1.4

2 4 107 94.8 1.4

6 112.8 100.0 1.2

3 4 101 89.5 1.1

6 104.6 92.7 0.8

4 4 109.9 97.4 1

6 110 97.5 1

5 4 111.6 98.9 0.7

6 114.7 101.7 0.8

6 4 109.1 96.7 1.7

6 111 98.4 1.3

7 4 110.9 98.3 1.9

6 110.7 98.1 0.9

8 4 117.2 103.9 1.2

6 113.4 100.5 1.4

9 4 111.4 98.7 1.1

6 115 101.9 0.8

10 4 111.7 99.0 1.9

6 110.3 97.8 0.9

γd ω
116

Table 40. TF-2 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 4 148.5 99.7 5.5

6 149.8 100.6 7.2

2 4 147 98.7 6.7

6 145.5 97.7 6

3 4 145.6 97.8 6.9

6 147.6 99.1 6.3

4 4 141.9 95.3 6.6

6 141.5 95.0 6.3

5 4 147.2 98.9 5.9

6 148.6 99.8 6.2

6 4 145.5 97.7 6.2

6 146.2 98.2 6.8

7 4 146.4 98.3 6.4

6 145.8 97.9 5.8

8 4 147.3 98.9 6

6 146.2 98.2 6.1

9 4

6

10 4

6

Blank cell = no data available.

γd ω
117

Table 41. TF-3 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 4 147.9 99.3 6.7

6 148.1 99.5 7

2 4 146.1 98.1 6.7

6 150.8 101.3 6.6

3 4 147.4 99.0 6.2

6 149 100.1 5.9

4 4 146.2 98.2 6.8

6 149.5 100.4 6.1

5 4 149.7 100.5 5.5

6 149.1 100.1 6.5

6 6 147.3 98.9 7.8

7 6 147.2 98.9 7.9

8 6 149.3 100.3 6.4

9 6 151.3 101.6 6.1

10 6 151.1 101.5 6.6

Table 42. TF-4 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor

ω

(percent)

1 4 149.2 100.2 6.3

2 6 147.6 99.1 6.3

3 6 147.7 99.2 6.2

4 6 147.9 99.3 7.3

5 6 149 100.1 7

6 6 149.7 100.5 7.1

7 6 149.4 100.3 6.8

8 6 151.1 101.5 6.2

9 6 152.6 102.5 5.6

10 6 151.1 101.5 6.8

γd ω

γd
118

Table 43. TF-5 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 148.1 99.5 4.9

2 6 148.1 99.5 6.1

3 6 150.2 100.9 6.5

4 6 151.7 101.9 6.3

5 6 152.7 102.6 5.7

6 6 148 99.4 6.5

7 6 151 101.4 5.7

8 6 151 101.4 5.7

9 6 148.7 99.9 6.1

10 6 154.5 103.8 5.4

Table 44. TF-6 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 148.9 100.0 6.4

2 6 147.9 99.3 6

3 6 150.1 100.8 6.5

4 6 151 101.4 6.5

5 6 149.9 100.7 6.7

6 6 149.6 100.5 6.5

7 6 150.3 100.9 6.9

8 6 151.9 102.0 5.8

9 6 150.7 101.2 5.5

10 6 148.9 100.0 6.2

γd ω

γd ω
119

Table 45. TF-7 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 150.1 100.8 6.3

2 6 151.2 101.5 5.9

3 6 150.1 100.8 6.2

4 6 149.1 100.1 6.1

5 6 152.3 102.3 6.1

6 6 150.2 100.9 5.8

7 6 151 101.4 5.9

8 6 148.9 100.0 6.3

9 6 149.5 100.4 5.9

10 6 152.5 102.4 5.9

Table 46. TF-8 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 149.1 100.1 7.1

2 6 148.7 99.9 6.9

3 6 151 101.4 6.4

4 6 150.6 101.1 6.3

5 6 151.5 101.7 6.5

6 6 150.5 101.1 6.5

7 6 151.2 101.5 6.3

8 6 151.1 101.5 6.4

9 2 147.6 99.1 6.2

10 2 146.9 98.7 6.5

10 2 150.8 101.3 5.9

γd ω

γd ω
120

Table 47. TF-9 Nuclear density test results.

Lift

Depth 

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 150.9 101.3 5.9

2 8 150.2 100.9 5.9

3 6 148.9 100.0 5.4

4 8 147.4 99.0 5.5

5 6 151.1 101.5 6.1

6 8 147 98.7 6.5

7 6 151.3 101.6 6.3

8 8 150.1 100.8 6.3

9 6 150.1 100.8 6.3

10 8 150.6 101.1 6

Table 48. TF-10 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 148.5 99.7 5.1

2 8 146.5 98.4 5.4

3 6 150.5 101.1 5.6

4 8 148.5 99.7 6.5

5 6 148.4 99.7 7.4

6 8 148.7 99.9 6.6

7 6 149.1 100.1 6.3

8 8 149.2 100.2 6.5

9 6 149.3 100.3 6.7

10 8 151 101.4 6.6

γd ω

γd ω
121

Table 49. TF-11 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 2 148 99.4 6.8

2 150.2 100.9 6.5

2 2 153.2 102.9 5.5

2 149.8 100.6 6.2

3 2 149.8 100.6 6.1

2 150.4 101.0 5.8

4 2 151.8 101.9 5.8

2 151.8 101.9 5.7

5 2 150.6 101.1 6.2

2 148.2 99.5 6.1

6 2 149.9 100.7 5.8

2 149.3 100.3 5.5

7 2 147.4 99.0 5.6

2 150.6 101.1 5.7

8 2 149.9 100.7 5.5

2 150 100.7 5.8

9 2 150.2 100.9 5.4

2 152.3 102.3 5.4

10 2 149.9 100.7 5.8

2 149.8 100.6 5.9

γd ω
122

Table 50. TF-12 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 2 149 100.1 6

2 148.7 99.9 5.8

2 2 148.1 99.5 5.7

2 149 100.1 6.3

3 2 150.3 100.9 6.5

2 148.2 99.5 6.6

4 2 149.3 100.3 6.4

2 148.5 99.7 6.0

5 2 147.6 99.1 5.8

2 150.4 101.0 5.9

6 2 148 99.4 6

2 148.8 99.9 5.5

7 2 148.5 99.7 6

2 152.2 102.2 5.6

8 2 147.9 99.3 6.1

2 148 99.4 6.4

9 2 151.1 101.5 6.2

2 149 100.1 6.1

10 2 149.9 100.7 5.0

2 151.1 101.5 5.0

γd ω
123

Table 51. TF-13 Nuclear density test results

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 4 148 99.4 6.8

4 146.6 98.5 6.3

4 147.4 99.0 5.7

2 4 147.6 99.1 5.8

4 147.8 99.3 5.7

4 150.9 101.3 5.7

3 6 151.9 102.0 6.8

4 150.5 101.1 6.6

4 6 149.9 100.7 6.3

4 147 98.8 5.9

5 8 146.8 98.6 6

4 150.6 101.1 6.4

6 6 152.4 102.4 6

4 148.2 99.50 5.8

7 6 149.4 100.3 5.9

4 153.2 102.9 5.7

Table 52. TF-14 Nuclear density test results.

Lift

Depth

(inch) (pcf)

Percent of 

Proctor (percent)

1 6 147.8 99.3 4.4

4 150.5 101.1 6.6

2 6 150.6 101.1 6.8

4 148 99.4 5.6

3 6 148.3 99.6 5.8

4 149.7 100.5 5.5

4 6 150.6 101.1 7

4 147.9 99.3 7.4

5 6 147.4 99.0 7.4

4 148.5 99.7 7

6 6 150 100.7 6.2

4 148 99.4 7.1

7 6 150.2 100.9 6.7

4 147.8 99.3 6.6

γd ω

γd ω

125

APPENDIX C. DEFORMATION INSTRUMENTATION LAYOUTS FOR PTS

Figure 107. Illustration. Instrumentation layout for DC tests and TF-1.

Figure 108. Illustration. Instrumentation layout for TF-2, TF-9.
126

Figure 109. Illustration. Instrumentation layout for TF-3, TF-4.

Figure 110. Illustration. Instrumentation layout for TF-5, TF-7.
127

Figure 111. Instrumentation layout for TF-6, TF-12.

Figure 112. Illustration. Instrumentation layout for TF-8.
128

Figure 113. Illustration. Instrumentation layout for TF-10.

Figure 114. Illustration. Instrumentation layout for TF-13
129

Figure 115. Illustration. Instrumentation layout for TF-14.

131

APPENDIX D. RAW DATA FOR PTS

Table 53. DC-1 PT Data.

Average 

Load

(lb/jack)

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00 0.00

6,000.29 2,666.80 0.33 0.43

7,691.24 3,418.33 0.43 0.56

10,891.56 4,840.69 0.58 0.76

13,882.17 6,169.85 0.82 1.08

16,743.42 7,441.52 1.02 1.34

18,617.17 8,274.30 1.19 1.56

20,737.31 9,216.58 1.35 1.78

23,679.35 10,524.15 1.56 2.05

26,749.52 11,888.68 1.85 2.43

29,760.19 13,226.75 2.12 2.79

32,796.45 14,576.20 2.60 3.43

36,725.03 16,322.23 2.99 3.93

38,843.79 17,263.90 3.24 4.26

41,800.05 18,577.80 3.57 4.70

45,218.53 20,097.12 3.82 5.02

49,169.75 21,853.22 4.20 5.52

52,488.78 23,328.34 6.04 7.95
132

Table 54. DC-2 PT Data.

Average 

Load

(lb/jack)

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00 0.00

5,787.51 2,572.23 0.59 0.78

7,200.47 3,200.21 0.75 0.99

8,940.54 3,973.57 0.92 1.21

10,952.92 4,867.97 1.17 1.53

12,974.76 5,766.56 1.40 1.83

15,051.84 6,689.71 1.64 2.15

17,077.36 7,589.94 1.88 2.46

19,204.93 8,535.52 2.17 2.84

21,211.57 9,427.36 2.50 3.28

25,264.01 11,228.45 3.19 4.18

27,203.24 12,090.33 3.50 4.59

29,368.95 13,052.87 3.73 4.89

31,354.88 13,935.50 4.01 5.26

34,309.37 15,248.61 4.47 5.86

37,455.51 16,646.89 4.76 6.24

40,418.89 17,963.95 4.83 6.34

43,298.49 19,243.77 4.90 6.43

51,134.47 22,726.43 5.37 7.04
133

Table 55. DC-3 PT Data.

Average 

Load

(lb/jack)

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00 0.00

1,649.95 733.31 0.24 0.31

4,305.23 1,913.44 0.52 0.68

6,976.55 3,100.69 0.78 1.03

9,692.94 4,307.98 1.03 1.35

12,450.76 5,533.67 1.29 1.70

14,973.40 6,654.85 1.54 2.02

17,582.22 7,814.32 1.83 2.41

20,165.77 8,962.56 2.12 2.78

23,249.20 10,332.98 2.42 3.19

26,217.45 11,652.20 2.77 3.64

29,407.36 13,069.94 3.08 4.05

32,442.93 14,419.08 3.37 4.43

35,518.85 15,786.16 3.70 4.87

38,565.19 17,140.08 4.04 5.32

41,538.90 18,461.73 4.42 5.82
134

Table 56. DC-4 PT Data.

Average 

Load

(lb/jack)

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00 0.00

2,724.22 1,210.76 0.16 0.22

5,017.17 2,229.85 0.38 0.51

7,568.27 3,363.68 0.59 0.77

9,949.51 4,422.01 0.82 1.08

12,971.35 5,765.05 1.11 1.46

15,926.47 7,078.43 1.40 1.84

19,025.69 8,455.86 1.70 2.24

21,821.66 9,698.52 1.98 2.61

24,883.53 11,059.35 2.26 2.97

27,993.61 12,441.61 2.58 3.40

30,894.62 13,730.94 2.94 3.87

34,053.34 15,134.82 3.30 4.34

36,906.96 16,403.09 3.54 4.66

39,975.59 17,766.93 3.91 5.15

42,919.44 19,075.31 4.25 5.60

45,841.51 20,374.00 4.52 5.95

47,205.64 20,980.28 4.70 6.18

50,450.81 22,422.58 4.95 6.52

55,229.31 24,546.36 5.36 7.06

60,188.19 26,750.31 5.81 7.64
135

Table 57. DC-5 PT Data.

Average 

Load

(lb/jack)

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00 0.00

622.79 276.79 0.37 0.48

2,739.19 1,217.42 0.68 0.89

4,857.51 2,158.89 0.96 1.26

7,021.37 3,120.61 1.20 1.57

9,060.95 4,027.09 1.43 1.88

11,088.41 4,928.18 1.68 2.21

14,307.26 6,358.78 2.02 2.66

17,440.25 7,751.22 2.41 3.17

20,525.62 9,122.50 2.83 3.73

23,676.51 10,522.90 3.26 4.29

26,657.83 11,847.92 3.84 5.05

29,865.77 13,273.68 4.24 5.58

32,794.82 14,575.47 4.64 6.10

35,897.19 15,954.31 5.05 6.64

43,500.00 19,333.33 5.92 7.79

48,500.00 21,555.56 7.92 10.42
136

Table 58. TF-1 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

369.70 0.08 0.10

556.35 0.08 0.11

978.69 0.16 0.21

1,842.65 0.33 0.43

2,804.93 0.55 0.72

3,696.98 0.75 0.99

4,646.86 0.99 1.30

5,578.45 1.22 1.59

6,465.19 1.45 1.90

7,448.68 1.70 2.22

8,340.86 1.96 2.57

9,232.90 2.24 2.94

10,164.52 2.51 3.29

11,057.92 2.80 3.67

11,952.08 3.08 4.04

12,761.14 3.28 4.31

13,472.62 3.86 5.07

14,465.32 4.76 6.24

15,231.20 4.99 6.54

16,000.57 5.31 6.96

16,994.27 5.55 7.27

17,862.85 6.13 8.04

18,781.38 7.36 9.66

20,486.98 8.31 10.90
137

Table 59. TF-2 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

829.67 0.03 0.03

1,760.89 0.09 0.12

2,604.89 0.17 0.22

3,559.11 0.25 0.33

4,509.78 0.35 0.46

5,438.67 0.46 0.60

6,372.00 0.57 0.75

8,220.00 0.83 1.09

10,084.00 1.13 1.48

11,006.67 1.29 1.70

13,812.44 1.92 2.53

16,591.11 2.64 3.48

19,472.89 4.22 5.55

22,352.00 6.47 8.51

25,278.67 8.71 11.46

Table 60. TF-3 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

883.61 0.03 0.04

2,692.89 0.25 0.33

3,590.22 0.41 0.54

5,333.33 1.01 1.33

6,311.11 1.22 1.60

8,088.89 1.98 2.60

9,866.67 2.67 3.50

10,755.56 3.13 4.10

13,333.33 4.42 5.80

14,666.67 5.41 7.10

16,156.89 7.52 9.87

17,503.56 10.52 13.80
138

Table 61. TF-4 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

1,496.46 0.04 0.06

2,157.57 0.08 0.10

3,211.69 0.27 0.36

4,128.22 0.47 0.62

5,072.21 0.70 0.92

6,015.01 0.94 1.23

6,930.25 1.16 1.53

8,837.16 1.71 2.25

10,703.07 2.34 3.08

14,250.89 3.34 4.40
139

Table 62. TF-5 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0 0 0

1,152 0.046214 0.060808

2,736 0.09353 0.123066

4,176 0.126134 0.165966

4,608 0.1733 0.228046

5,472 0.289825 0.381349

6,480 0.455363 0.599162

8,352 0.892178 1.173918

10,080 1.466227 1.929246

11,088 1.761062 2.317187

13,968 2.954815 3.887914

16,704 4.390352 5.776779

19,584 5.947989 7.826301

22,464 8.176841 10.759

25,344 9.917697 13.0496

10,944 9.843107 12.95146

0 8.705762 11.45495

0 8.700209 11.44764

2,448 0.512515 12.12201

3,744 0.663577 12.32077

5,616 0.822237 12.52953

8,496 1.007906 12.77384

11,232 1.19151 13.01542

16,992 1.785724 13.79728

22,464 3.312276 15.8059

25,920 4.924452 17.92719
140

Table 63. TF-6 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

889.74 0.06 0.08

1,764.73 0.20 0.27

2,522.17 0.26 0.34

3,718.49 0.38 0.50

4,546.27 0.48 0.64

5,357.84 0.65 0.85

6,255.09 0.73 0.96

8,011.64 0.93 1.22

9,850.08 1.82 2.39

13,449.91 2.27 2.99

18,706.79 3.16 4.15

21,447.16 3.68 4.84

24,041.44 4.24 5.58

26,687.14 4.76 6.27

29,321.85 5.34 7.02

30,643.85 5.79 7.62

31,937.57 6.25 8.22

33,305.27 6.83 8.99

35,809.33 7.66 10.08

38,714.79 8.76 11.52

41,250.78 10.75 14.15

43,827.56 11.95 15.73
141

Table 64. TF-7 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0 0 0

906.5827 0.14 0.19

1,843.848 0.31 0.40

2,727.728 0.48 0.63

3,579.004 0.65 0.86

4,443.513 0.90 1.19

5,359.769 1.08 1.42

6,237.56 1.29 1.70

7,992.877 1.75 2.30

9,808.05 2.29 3.01

10,739.62 2.59 3.41

13,366.63 3.65 4.80

15,953.95 4.89 6.44

18,656.81 6.54 8.61

21,257.23 7.80 10.27

24,085.41 9.54 12.55

26,573.67 12.46 16.39
142

Table 65. TF-8 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

901.25 0.19 0.24

1,784.19 0.39 0.51

2,673.33 0.60 0.78

3,610.08 0.81 1.07

4,460.27 1.00 1.32

5,317.64 1.20 1.58

6,194.72 1.42 1.87

7,942.55 1.87 2.46

9,762.13 2.38 3.13

10,636.89 2.66 3.50

13,345.60 3.62 4.76

16,003.43 4.70 6.18

18,547.61 5.90 7.76

21,274.89 7.35 9.67

23,999.31 8.62 11.34

26,608.37 10.10 13.29

29,178.74 13.51 17.77
143

Table 66. TF-9 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

434.85 0.06 0.08

573.32 0.07 0.09

746.42 0.09 0.12

1,225.31 0.14 0.19

1,898.80 0.24 0.31

2,541.19 0.32 0.42

3,146.98 0.41 0.54

3,725.78 0.50 0.65

4,334.89 0.63 0.84

5,585.85 0.85 1.12

6,790.22 1.08 1.43

7,433.04 1.24 1.63

9,292.90 1.74 2.29

11,195.61 2.46 3.24

13,072.80 3.36 4.43

14,953.66 5.17 6.80

16,806.30 6.49 8.53

18,680.80 8.05 10.59

20,537.79 9.95 13.09

22,360.24 11.83 15.56
144

Table 67. TF-10 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

375.25 0.09 0.12

1,115.40 0.28 0.37

1,443.58 0.35 0.46

1,641.11 0.39 0.51

2,060.36 0.50 0.66

2,478.58 0.64 0.84

2,915.38 0.77 1.02

3,691.10 1.07 1.40

4,536.46 1.46 1.92

4,958.59 1.68 2.22

6,186.24 2.53 3.33

7,432.09 3.86 5.08

8,625.99 6.32 8.32

8,838.17 9.05 11.91

10,339.14 10.85 14.28
145

Table 68. TF-11 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

1,321.94 0.08 0.10

1,712.69 0.21 0.27

2,676.26 0.35 0.46

3,506.48 0.50 0.66

4,368.60 0.67 0.89

5,242.11 0.87 1.15

6,148.03 1.08 1.43

7,785.80 1.56 2.05

9,525.78 2.15 2.83

10,468.16 2.54 3.34

13,028.22 3.65 4.81

15,745.28 5.04 6.63

18,444.67 7.03 9.25

21,111.16 8.68 11.43

23,276.92 9.72 12.79
146

Table 69. TF-12 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

995.38 0.11 0.14

2,018.60 0.20 0.26

3,079.10 0.29 0.39

4,185.50 0.39 0.52

5,360.08 0.51 0.67

6,463.52 0.63 0.82

7,552.20 0.75 0.98

9,702.94 1.02 1.34

11,883.24 1.32 1.74

12,969.50 1.54 2.03

16,219.10 2.27 2.99

19,473.98 4.39 5.77

22,908.22 7.05 9.28

26,174.57 9.09 11.96

29,062.81 10.16 13.37
147

Table 70. TF-13 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0 0 0

762.3172 0.098888 0.125573

1,333.227 0.162839 0.20678

1,879.568 0.26315 0.334158

2,320.298 0.38099 0.483797

3,020.67 0.511435 0.649442

3,600.98 0.668502 0.848892

3,931.689 0.836782 1.06258

5,394.454 1.27596 1.620266

6,626.904 1.847268 2.345737

7,227.084 2.249614 2.856652

9,002.369 4.368862 5.547761

10,789.2 6.084915 7.726876

12,984.6 9.703566 12.32199
148

Table 71. TF-14 PT Data.

Applied 

Pressure

(psf)

Vertical 

Settlement

(inch)

Vertical 

strain

(percent)

0.00 0.00 0.00

580.06 0.16 0.21

1,191.94 0.25 0.32

1,829.39 0.35 0.46

2,466.98 0.45 0.59

3,001.13 0.54 0.71

3,611.30 0.64 0.84

4,225.24 0.74 0.98

5,402.42 0.94 1.24

6,620.61 1.15 1.51

7,233.56 1.28 1.68

9,042.58 1.65 2.18

10,965.38 2.07 2.72

12,738.15 2.53 3.33

14,571.33 3.08 4.05

16,460.84 3.68 4.84

18,314.54 4.45 5.85

20,140.13 5.51 7.25

22,071.79 7.13 9.39

23,603.11 9.65 12.69
149

ACKNOWLEGEMENTS

The authors would like to acknowledge the hard work and effort of Ms. Jan Li, Mr. Moustafa 

Ibrahim-Awad, Mr. Tim Tuggle, Mr. Jeffrey Lin, and Mr. Daniel Powers in helping to construct 

and test the performance tests at Turner-Fairbank Highway Research Center. In addition, thanks 

to Dr. Fassil Beshah for coordinating use of the Structures Laboratory. For the tests conducted in 

Defiance County, OH, the authors wish to thank Mr. Warren Schlatter and Mr. Jeff Timbrook, 

along with the participants of each GRS EDC Validation Session. Finally, the authors would also 

like to extend their appreciation for the technical review comments prepared by (in alphabetical 

order) Mr. Daniel Alzamora of the Resource Center, Khalid Mohamed of the Office of Bridge 

Technology, and Dr. Justin Ocel of the Office of Infrastructure Research and Development.

151

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