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Creation of a Laboratory Testing Device to Evaluate Instability Rutting in Asphalt Pavement

Permanent Link: http://ufdc.ufl.edu/UFE0021256/00001

Material Information

Title: Creation of a Laboratory Testing Device to Evaluate Instability Rutting in Asphalt Pavement
Physical Description: 1 online resource (176 p.)
Language: english
Creator: Novak, Marc E
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2007

Subjects

Subjects / Keywords: Civil and Coastal Engineering -- Dissertations, Academic -- UF
Genre: Civil Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Near-surface rutting within the asphalt layer, known as instability rutting, has become a costly mode of pavement distress on today's roads. Instability rutting reduces ridability, increases the potential for ponding leading to the potential for hydroplaning, and necessitates costly rehabilitation. This research sought to identify the mechanisms behind instability rutting and to development a laboratory testing device that could evaluate an asphalt mixtures's ability to resist instability rutting. Three-dimensional FEM analyses using the program ADINA was used to identify stresses radial tires induce to the pavement. The analyses indicated that radial tires induce high shear stress at low confinements within the asphalt pavement in areas where instability rutting is observed to occur at levels not predicted by traditional uniform vertical loading. High shear at low confinement is believed to be a key factor in the mechanism of instability rutting. Based on the results of the FEM analyses, a laboratory device that could replicate the high shear stresses at low confinements was sought. A laboratory device that could achieve this was the hollow cylinder testing device. By testing asphalt specimens under the critical stress condition of high shear at low confinement, it is believed will help in evaluating a mixture's ability to resist instability rutting. A hollow cylinder testing device was developed at the University of Florida within the frame work of existing laboratory equipment. The hollow cylinder device was developed such that axial, shear, and confinement stresses (inner and outer pressures) could be applied simultaneously to induce stress states similar to those identified by the FEM. Laboratory prepared asphalt specimens of known instability rutting performance were tested with the hollow cylinder device under cyclic stress critical stress and axial and shear strains were measured. The hollow cylinder is a device that will help lead to a possible screening tool to determine a mixture's susceptibility to instability rutting and will lead to insight into the mechanisms behind instability rutting.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Marc E Novak.
Thesis: Thesis (Ph.D.)--University of Florida, 2007.
Local: Adviser: Birgisson, Bjorn.
Local: Co-adviser: Roque, Reynaldo.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2007
System ID: UFE0021256:00001

Permanent Link: http://ufdc.ufl.edu/UFE0021256/00001

Material Information

Title: Creation of a Laboratory Testing Device to Evaluate Instability Rutting in Asphalt Pavement
Physical Description: 1 online resource (176 p.)
Language: english
Creator: Novak, Marc E
Publisher: University of Florida
Place of Publication: Gainesville, Fla.
Publication Date: 2007

Subjects

Subjects / Keywords: Civil and Coastal Engineering -- Dissertations, Academic -- UF
Genre: Civil Engineering thesis, Ph.D.
bibliography   ( marcgt )
theses   ( marcgt )
government publication (state, provincial, terriorial, dependent)   ( marcgt )
born-digital   ( sobekcm )
Electronic Thesis or Dissertation

Notes

Abstract: Near-surface rutting within the asphalt layer, known as instability rutting, has become a costly mode of pavement distress on today's roads. Instability rutting reduces ridability, increases the potential for ponding leading to the potential for hydroplaning, and necessitates costly rehabilitation. This research sought to identify the mechanisms behind instability rutting and to development a laboratory testing device that could evaluate an asphalt mixtures's ability to resist instability rutting. Three-dimensional FEM analyses using the program ADINA was used to identify stresses radial tires induce to the pavement. The analyses indicated that radial tires induce high shear stress at low confinements within the asphalt pavement in areas where instability rutting is observed to occur at levels not predicted by traditional uniform vertical loading. High shear at low confinement is believed to be a key factor in the mechanism of instability rutting. Based on the results of the FEM analyses, a laboratory device that could replicate the high shear stresses at low confinements was sought. A laboratory device that could achieve this was the hollow cylinder testing device. By testing asphalt specimens under the critical stress condition of high shear at low confinement, it is believed will help in evaluating a mixture's ability to resist instability rutting. A hollow cylinder testing device was developed at the University of Florida within the frame work of existing laboratory equipment. The hollow cylinder device was developed such that axial, shear, and confinement stresses (inner and outer pressures) could be applied simultaneously to induce stress states similar to those identified by the FEM. Laboratory prepared asphalt specimens of known instability rutting performance were tested with the hollow cylinder device under cyclic stress critical stress and axial and shear strains were measured. The hollow cylinder is a device that will help lead to a possible screening tool to determine a mixture's susceptibility to instability rutting and will lead to insight into the mechanisms behind instability rutting.
General Note: In the series University of Florida Digital Collections.
General Note: Includes vita.
Bibliography: Includes bibliographical references.
Source of Description: Description based on online resource; title from PDF title page.
Source of Description: This bibliographic record is available under the Creative Commons CC0 public domain dedication. The University of Florida Libraries, as creator of this bibliographic record, has waived all rights to it worldwide under copyright law, including all related and neighboring rights, to the extent allowed by law.
Statement of Responsibility: by Marc E Novak.
Thesis: Thesis (Ph.D.)--University of Florida, 2007.
Local: Adviser: Birgisson, Bjorn.
Local: Co-adviser: Roque, Reynaldo.

Record Information

Source Institution: UFRGP
Rights Management: Applicable rights reserved.
Classification: lcc - LD1780 2007
System ID: UFE0021256:00001


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CREATION OF A LABORATORY TESTING DEVICE TO EVALUATE INSTABILITY
RUTTING IN ASPHALT PAVEMENTS



















By

MARC E. NOVAK


A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL
OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT
OF THE REQUIREMENTS FOR THE DEGREE OF
DOCTOR OF PHILOSOPHY

UNIVERSITY OF FLORIDA

2007



































2007 Marc E. Novak

































To my wife Patricia









ACKNOWLEDGMENTS

I would like to acknowledge those individuals who were involved in the advancement of

this research. First, I would like to thank my advisor and mentor Dr. Bjorn Birgisson for his time

and support and the advice given to me. Acknowledgments also must go to my committee

members for their time and input on numerous aspects of the research Dr. Reynaldo Roque, Dr.

Michael McVay, and Dr. Joe Tedesco. I would also like to thank my outside committee

member, Dr. Bhavani Sankar, for his input.

Special thanks and appreciation go to the University of Florida staff. I would like to

especially thank Mr. George Lopp for his time and expertise and graduate students including

Aditya Ayithi, Linh Pham, Dinh Nguyen and Christos Drakos for the numerous discussions and

debates on research, but more importantly, their friendship. Special thanks also to Jason

Crockett and William Vash, their friendship, support, and laughter is greatly appreciated. In

addition, I would like to thank my father and mother-in-law, John and Rubi Kwong, for their

support and encouragement.

I greatly appreciate the care and support of the St. Augustine Catholic Church community.

I also appreciate the time and understanding afforded to me by Luis Mahiquez, Henri Jean,

Jeanne Berg, and Larry Moore of Tierra, Inc.

I would also like to thank my parents, John and Elaine Novak, for being a source of

support throughout my studies. In addition, I appreciate the love and support of my grand-parents

Michael and Helen Novak, Edmond Wojtowicz, Eleanor Wojtowicz Buber, and my step-

grandfather, Michael Buber. Finally, and most importantly, I would like to thank my wife,

Patricia; it was her love, patience, and support that helped me the most.










TABLE OF CONTENTS

page

A C K N O W L E D G M E N T S ..................................................................... ......... ........................... 4

L IST O F T A B L E S ....................... ................................................................. . 8

L IST O F FIG U R E S ............................................................................... 9

ABSTRACT ........................................... .. ......... ........... 14

CHAPTER

1 INTRODUCTION ............... .......................................................... 16

B ack g rou n d ................... ...................1...................6..........
Problem Statement ................................................................ ..... ..... ........ 18
H y p o th e sis ..........................................................................1 8
O bje ctiv e s ................... ...................1...................9..........
S cop e ................ .............................................................. 2 0
Research Approach ................................................................................20

2 L IT E R A T U R E R E V IE W ................................................................................................. 2 1

P erm anent D reform action ................................................................................ 2 1
T ire-Interface Stresses ................................................................23
Accelerated Pavement Testers ................................. ........................... .........27
T torture T est D devices ............................................................................................... ........29
L ab o ra to ry T e sts ............................................................................................................... 3 0
C reep T e sts ...........................................................3 0
C om plex M odulus Testing ...........................................................32
T riax ia l T e stin g ............................................................................................................... 3 4
Superpave Shear Tester .................................................................. .. ......... 36
H follow C y lin der T testing ........................................................................................... 3 8
S u m m ary ................... ...................4...................4..........

3 HEAVY VEHICLE SIMULATOR TESTING ON ASPHALT PAVEMENTS ...................46

Instability Rutting in Modified and Unmodified Pavements Under HVS Loading ............47
Rutting Depths .........................................................................47
Rut Propagation Analysis through Surface Profiling .................................................51
Volumetric Analysis of Asphalt Cores ......... ......... ............ .......... 55
One-way Versus Two-way Directional Loading .............. ............. ........ 59
Quantification of Shear Stresses Within Asphalt Pavement ................................ ............. 64
A analytical A approach ......... .. ......... ...............................................................................70
Sum m ary and C conclusions ....................................................................................................75









4 THE APPROXIMATION OF NEAR-SURFACE STRESS STATES IN ASPHALT
CONCRETE THROUGH THREE-DIMENSIONAL FINITE ELEMENT ANALYSES ....76

B ack g rou n d ........................... ........................................................................................... 7 6
Three-dim ensional Finite Elem ent M odel ........................................ ......................... 77
Pavement Structure and Loading Conditions Analyzed .....................................................80
P avem ent Structure........ ...................................................................... ......... .......80
Loading Conditions A analyzed ............................................................ ............... 80
Load Application ...................................................................... ......... 83
Axisym m etric m odel ................................... .. ... .......... ...... ........ .. 85
Three-D im ensional Solution Process ........................................ ......................... 87
Results of Three-D im ensional Analysis ...................................................... ..... .......... 88
Significance of High Shear at Low Confinement................................................................ 97
Sum m ary and C onclu sions .......................................................................... .....................99

5 CREATION OF A HOLLOW CYLINDER DEVICE............... ................. 101

In tro d u ctio n .................. ...................................... .................. ................ 10 1
Hollow Cylinder Specimen Dimensions ................... ....... ...............102
Closed-form analysis of stresses across hollow cylinder wall ............................104
Finite element analysis of stress non-linearity and concentrations ...................110
Creation of a Hollow Asphalt Concrete Specimen ....................... ...............................115
H ollow Cylinder Specim en Production......... ............................................ .................. 116
Aggregate Preparation and Batching............................................................. ....... 116
M ix in g .............. ... ............... ................ ......................... 1 17
Short Term Oven Aging (STOA) and Compacting......................................................117
C o rin g ...................................................................................................................... 1 18
H follow Cylinder Testing D vice .......................................................... ............... 119
S u m m ary ................... ...................1...................2.........6

6 MATERIAL SELECTION AND TESTING PROGRAM............................127

Intro du action ............. ...... ......... ...................27..........
Material Selection.......... .................... .. .... .... .... ........... 127
Testing Program ................................................. 129

7 COMPLEX MODULUS AND EXTENSION TESTING................... ...............131

C om plex M odulu s T testing ........................................................................ ..................... 13 1
B background ............. .. ....... ......... ......................................13 1
T testing P program ....................................................... 135
E and G T testing R results .................................... ........... .............. .........................136
E v alu atio n s ................................. ....................... ................................. 14 0
E xten sion T testing ...............................................................14 1
Sum m ary and C onclu sions ........................................ .......... ................. ..........................144





6









8 TORSIONAL EXTENSION TESTING................................................................... 145

In tro d u ctio n .................. .................................................................................................. 14 5
L laboratory L oading C onditions............................................. ......................................... 146
L laboratory T testing P program ........................................................................ ................... 148
L laboratory T testing R esults........................................................................ ......... ........... 148
Testing to Failure ..................................................................148
Load Relaxation at Tertiary Range, then Load Re-Application..................................152
Load Relaxation During Secondary Phase and Re-Loading ............... ...............155
E v alu atio n s ........................................................................16 0
C onclusions.....................................................................162

9 CON CLU SION .......... ............................................ ... .. .. .................. .. 163

Summary of Work ................................. .. .. ..... ...... ............ 163
F future W ork ......................................................164

L IST O F R E F E R E N C E S ............................................................................. ..........................165

B IO G R A PH IC A L SK ETCH ............................................................................... .......... ........ 176


































7









LIST OF TABLES


Table page

3-1 T est sections in H V S testing ...................................................................... ..................48

3-2 Summary of asphalt cores from pavement sections 5A and 1B post-testing...................56

3-3 Summary of core data through use of phase diagrams. .................................................58

3-4 Deformation of an element in an asphalt pavement near the edge of the tire in the
Y Z -plane (transverse plane) ....................................................................... ..................63

3-5 Properties of pavement layers used in the finite element analysis..................................66

3-6 Values used in Burgers model to represent asphalt concrete pavement..........................74

3-7 Results of Burger model under two-way and one-way loading..................... ........ 74

4-1 Material properties of the various layers used in the analysis. ........................ .........80

5-1 Possible hollow cylinder specimen dimensions.................................... ............... 104

5-2 Inner and outer pressures applied to proposed hollow cylinder geometries..................105

5-3 Beta 3 shear stress non-uniformity quantities for the various geometries.....................107

6-1 Superpave mixtures employed for testing program...................................................... 128

6-2 Tests performed with the hollow cylinder test device. .................................................129

7-1 Summary of air void contents of different specimen geometries .................................. 137

8-1 Testing matrix for cyclic torsional testing. ........................................... ............... 148

8-2 Three ranges of the axial and shear strain response of the mixtures ............................151









LIST OF FIGURES


Figure page

2-1 Structural or consolidation rutting .............................................................................. 22

2-2 Instability rutting..................... ....... ..... .. .... .... ...............23

2-3 V ertical stress distribution of a radial tire................................. ........................ .. ......... 24

2-4 Lateral stress distribution of a radial tire ........................................ ....................... 25

2-5 The Smithers Scientific Inc. tire contact stress measurement device...............................25

2-6 SST test cham ber. ..................................... .. ......... ......... .... 37

2-7 Plan view of a hollow cylinder with outer radius, ro, and inner radius, ri, and subject
to varying outer (Po) and inner pressure (Pi)............................................. ............... 41

2-8 Profile view of a hollow cylinder specimen. ........................................ ............... 41

3-1 The Mark IV Heavy Vehicle Simulator device. ..................................... ............... 46

3-2 Rut depth progression for the various mixes. ....................................... ............... 49

3-3 Trench cut on one of the sections tested during the HVS study. ............. ..................50

3-4 Transverse cross-section of the asphalt pavement after testing............... ... ...............50

3-5 Average transverse profile created by averaging the 58 longitudinal measurements
from the laser profiler. ............................................. .. .. ......................51

3-6 The 59 longitudinal values averaged. ........................................ .......................... 52

3-7 Average transverse profile with ratio equation....................... .......... ... ............ 53

3-8 The ratio of area elevated to area depressed for all 500 C sections. ..................................54

3-9 Final transverse rut profile of sections 1B and 5A. ................................... ..................... 55

3-10 Phase diagram of the asphalt cores. .........................................................................57

3-11 Comparison of one-way and two-way directional loading with no tire wander................60

3-12 Comparison of one-way and two-way directional loading with tire wander...................60

3-13 Instability rutting with directional axes. ........................................ ........................ 61

3-14 Two-way versus one-way directional loading. ...................................... ............... 62









3-15 Deformation of an element in an asphalt pavement near the edge of the tire in the
X Z -plane (longitudinal plane)......................................... .............................................64

3-16 Area of pavement system where shear stresses were obtained............... ............... 67

3-17 Shear stress in Y Z -plane. ......................................................................... .....................68

3-18 Shear stress in X Z -plane. ......................................................................... .....................68

3-19 Shear stress in X Y -plane ......... ......................................................... ............................69

3-20 Burgers m odel............... ......... ......................................................... .. 70

3-21 Longitudinal shear stress pattern in one-way loading............................. ...............73

3-22 Longitudinal shear stress pattern in two-way loading. ................... ............................. 73

4-1 Three-dimensional finite element mesh used in the analysis............... ...............79

4-2 Plan view of the three-dimensional finite element mesh used in the analysis ................79

4-3 Variation in vertical stress across the 5-rib radial tire used in the analysis.....................81

4-4 Variation in vertical stress across 4-rib radial tire used in the analysis...........................82

4-5 Variation in horizontal stress across 5-rib radial tire used in the analysis......................82

4-6 Variation in horizontal stress across 4-rib radial tire used in the analysis......................83

4-7 Contact area and the radial tire nodal forces used in the pavement response analysis......85

4-8 Cross-sectional view of applied tire nodal forces used in the pavement response
an aly sis.......... ........................................................... .......................... ............... 8 5

4-9 A cross-sectional view of the axisymmetric finite element mesh used for comparison
purposes ......................................................... .................................86

4-10 Comparison of shear stresses under the edge of a two-dimensional axisymmetric
circular uniform vertical load predicted with BISAR and ADINA. .................................87

4-11 Location within asphalt pavement from finite element analysis for preliminary stress
state ev alu atio n ...................................................... ................ 8 9

4-12 Maximum shear stress in longitudinal direction along outer tire rib...............................90

4-13 Confinement stress in longitudinal direction along outer tire rib. ....................................91

4-14 Confinement and maximum shear stress path. ...................................... ............... 91









4-17 Maximum shear stress magnitude (in kPa) and direction under the 5-rib radial tire
loading condition. .......................................... ............................ 94

4-18 Maximum shear stress magnitude (in kPa) and direction under the 4-rib radial tire
loading condition. .......................................... ............................ 95

4-19 Maximum shear stress magnitude (in kPa) and direction under the uniform loading
condition ........................................................ .................................95

4-21 Stress states at point A for the three loading conditions analyzed............................... 96

4-22 Stress state at point B for the three loading conditions analyzed. ....................................97

4-23 Typical strength envelopes in hot mix asphalt and granular materials............................98

5-1 Ratio of radial stress to average radial stress across the wall of the specimen for
various geom entries. ........................ ......... .. ..... ..... .. ............106

5-2 Ratio of tangential stress to average tangential stress across the wall of the specimen
for various s geom etries........ .................................................................... .......... ....... 106

5-3 Ratio of shear stress to average shear stress across the wall of the specimen for
various geom entries. ........................ ......... .. ..... ..... .. ............107

5-4 Hollow cylinder specimen dimensions. .........................................................................109

5-5 Three-dimensional finite element model of the hollow cylinder specimen for end
effects analysis. .............. ............................... .......... ......... ......... 110

5-6 Tangential (hoop) stress distribution across hollow cylinder wall from finite element
an aly sis ................. ........................ .. .. ......... .......................................... 1 1 1

5-7 Radial stress distribution across hollow cylinder wall from finite element analysis.......112

5-8 Three vertical locations where stress distributions across the wall from the finite
element solution were compared to the closed-form solution. ........... ...............113

5-9 Comparison of the tangential stress across the hollow cylinder wall from the closed-
form solution and the finite element analysis with end constraints..............................113

5-10 Comparison of the radial stress across the hollow cylinder wall from the closed-form
solution and the finite element analysis with end constraints............... ... ............ 114

5-11 Hollow cylinder specimen production......... ...................... .......................116

5-12 Coring process to obtain hollow cylinder specimen ...................................................118

5-13 Plan view of the pedestal. ..................................................................... ..................... 120









5-14 Side view of pedestal. .................. ............... ................ .......... ...... ........ .. 121

5-15 Plan view of the bottom end cap .............................................. .. ............................ 122

5-16 Side view of the bottom end cap .............................................. .. ............................ 122

5-18 Side view of the top end cap. ...................................................................... 123

5-19 Plan view of the connection cap. .............................................................................. 124

5-20 Side view of the connection cap. .............................................................................. 124

5-21 The hollow cylinder testing device ................................. ............... ............... 125

5-22 Hollow cylinder testing device inside plexiglass chamber and surrounded by latex
m em brane........................................................ ................................ 126

7-1 Torsional shear test for hollow cylinder column. ................................. .................132

7-2 E* versus microstrain for different WRC1 specimen geometries under axial loading
at 1H Z ......................................................... ..................................137

7-3 G* versus microstrain for different WRC1 specimen geometries under torsional
loading at 1H Z ......................................................................... 138

7-4 E* versus microstrain for different SP1 specimen geometries under axial loading at
1H Z .. .......................................................................................... 13 8

7-5 G* versus microstrain for different SP1 specimen geometries under torsional loading
at 1H Z ......................................................... ..................................139

7-6 WRC1 and SP1 hollow cylinder specimen E* results versus microstrain.....................140

7-7 WRC1 and SP1 specimen G* results versus microstrain. .............................................140

7-8 P-Q diagram of stress path used for extension testing..............................................142

7-9 Axial strain versus deviator stress, q, for different WRC 1 specimen geometries ..........142

7-10 Axial strain versus deviator stress, q, for different SP1 specimen geometries..............143

8-1 Stress state imposed on hollow cylinder samples during testing................................... 146

8-2 Cyclic stress path imposed on asphalt specimens for testing. .......................................147

8-3 Shear strain versus number of cycles of two SP1 and WRC1 specimens under axial-
torsional-extension testing ................................................................... ..................... 149









8-4 Axial strain versus number of cycles of two SP 1 and WRC 1 specimens under axial-
torsional-extension testing ................................................................... ..................... 150

8-5 Angle of axial to shear strain,Y of two SP1 and WRC1 specimens under axial-
torsional-extension testing. ..................................................................... ...................150

8-6 Shear strain versus cumulative number of cycles for WRC1 and SP1 specimens. .........153

8-7 Axial strain versus cumulative number of cycles. .......................... ........................... 153

8-8 Angle Y versus cumulative number of cycles. .......................................... 154

8-9 Shear strain versus tim e. ........................................................................ ....................154

8-10 Axial stain versus time ...... .............. ...... .. ....... ................. 155

8-12 Axial strain versus cumulative number of cycles. ................................ .....................157

8-14 Shear strain versus number of cycles ............. .... ....................................... 158

8-15 Axial strain versus number of cycles. ................ ................................. ....... .... 159

8-16 Angle versus number of cycles. ...... ...................................................................... 159









Abstract of Dissertation Presented to the Graduate School
of the University of Florida in Partial Fulfillment of the
Requirements for the Degree of Doctor of Philosophy

CREATION OF A LABORATORY TESTING DEVICE TO EVALUATE INSTABILITY
RUTTING IN ASPHALT PAVEMENTS

By

Marc E. Novak

August 2007

Chair: Bjorn Birgisson
Major: Civil Engineering

Near-surface rutting within the asphalt layer, known as instability rutting, has become a

costly mode of pavement distress on today's roads. Instability rutting reduces ridability,

increases the potential for ponding leading to the potential for hydroplaning, and necessitates

costly rehabilitation. This research sought to identify the mechanisms behind instability rutting

and to development a laboratory testing device that could evaluate an asphalt mixture's ability to

resist instability rutting.

Three-dimensional finite element analysis using the program ADINA was used to identify

stresses radial tires induce to the pavement. The analyses indicated that radial tires induce high

shear stress at low confinements within the asphalt pavement in areas where instability rutting is

observed to occur at levels not predicted by traditional uniform vertical loading. High shear at

low confinement is believed to be a key factor in the mechanism of instability rutting.

Based on the results of the finite element analyses, a laboratory device that could replicate

the high shear stresses at low confinements was sought. A laboratory device that could achieve

this was the hollow cylinder testing device. By testing asphalt specimens under the critical stress









condition of high shear at low confinement, it is believed will help in evaluating a mixture's

ability to resist instability rutting.

A hollow cylinder testing device was developed at the University of Florida within the

frame work of existing laboratory equipment. The hollow cylinder device was developed such

that axial, shear, and confinement stresses (inner and outer pressures) could be applied

simultaneously to induce stress states similar to those identified by the finite element analyses.

Laboratory prepared asphalt specimens of known instability rutting performance were tested

with the hollow cylinder device under cyclic stress critical stress and axial and shear strains were

measured.

The hollow cylinder is a device that will help lead to a possible screening tool to determine

a mixture's susceptibility to instability rutting and will lead to insight into the mechanisms

behind instability rutting.









CHAPTER 1
INTRODUCTION

Background

A major distress mode in flexible (asphalt concrete) pavements is rutting. Rutting is the

mechanism that produces depressions in the wheel-paths of asphalt concrete pavements. Rutting

is the result of volumetric compression and/or shear deformation of one or more layers of the

pavement system (asphalt concrete, base, and/or subgrade) under repeated traffic loadings.

Rutting reduces serviceability and creates the potential for hydroplaning due to the accumulation

of water in the wheel-path ruts.

One form of rutting is known as "instability rutting." Instability rutting is rutting which is

confined only to the asphalt concrete layer. Instability rutting in asphalt pavements is primarily

due to the lateral displacement of material within the asphalt concrete layer. Instability rutting is

generally seen in pavements with a thick asphalt concrete layer (high trafficked roadways) and is

the predominant mode of premature failure is modem asphalt pavements.

Instability rutting is attributed strictly to the asphalt mixture properties and usually occurs

within the top 2-3 inches of the asphalt concrete layer. Instability rutting occurs when the

structural properties of the compacted pavement are inadequate to resist the stresses imposed

upon it. Despite instability rutting being the predominant mode of premature rutting failures in

modern flexible pavements, current pavement structural design approaches do not deal with

rutting in the asphalt concrete layer.

In 1987, Congress established the Strategic Highway Research Program (SHRP) with the

objective of improving the performance and durability of roadways in the United States. The

Superpave (Superior Performing Asphalt Pavements) mix design method was the result of the

SHRP to create a rational mix design method that could minimize distress in asphalt concrete









pavements, including rutting. The Superpave mixture design, although an improvement on

previous mixture design methods, is based solely on volumetrics. In Superpave mix design, an

asphalt mixture's susceptibility to instability rutting is mitigated by ensuring proper volumetric

quantities of the asphalt mixture are satisfied. Superpave has yet to incorporate a test that

directly measures or evaluates a mixture's resistance to rutting. The desire for a simple and direct

quantitative and/or qualitative measure of a mixture's susceptibility to instability rutting is

currently being sought in conjunction with the current Superpave mix design.

Today, most of the intra-continental commerce in the United States is transported through

trucks as opposed to rail in years past. In 1988, 2.2 million miles of our nation's roadways had

asphalt concrete surfaces with 91% of 2 trillion annual vehicle miles occurring on these asphalt

concrete roadways [Federal Highway Administration (FHWA) 1988]. Since that time, this

number has increased.

Over the last 30 years, the type of tires trucks are using is changing as well. Today, more

than 98% of trucks use radial tires, as opposed to bias-ply in years past. Radial tires allow for

higher inflation pressures and thus, the ability to carry higher loads. Studies have shown that

radial tires impart quantitatively and qualitatively different stresses to asphalt pavements [Myers

1997]. It has been shown that radial tires are more detrimental to the pavement's surface than

previous bias-ply tires [Myers et al. 1999].

Numerous performance prediction tests have been presented, discussed, and implemented

by different transportation agencies in order to determine an asphalt mixture's susceptibility to

instability rutting. In general, prediction tests may either be considered physical or index.

Physical tests can be simply defined as "torture tests". Torture tests subject an asphalt concrete

pavement or specimen to loading conditions that mimic field conditions. Large scale torture









testing is done on actual pavement sections through the use of vehicle simulators. Small scale

torture testing is done with loaded wheel testers (LWTs). LWTs subject a laboratory created

asphalt specimen to repeated miniature moving wheel loadings. The measured response (degree

of deformation) is then linked to field performance (rutting). Some common LWTs include the

Hamburg Wheel-Tracking Device, the French Pavement Rutting Tester, and the Asphalt

Pavement Analyzer.

Index tests are generally laboratory tests that attempt to determine one or more

fundamental material properties of asphalt concrete or the response of asphalt concrete under

certain loading conditions and link that property or response to overall rutting performance. Such

laboratory index tests that are concurrently being employed are creep (cyclic and static), complex

modulus, triaxial, and hollow cylinder tests. Unlike LWTs, laboratory tests place an equivalent

stress state throughout the specimen and can measure fundamental properties of the asphalt

mixture. Both approaches have merits and deficiencies and are currently being evaluated.

Problem Statement

Factors that influence or contribute to the amount of rutting and/or an asphalt concrete

pavement's resistance to rutting have not been clearly identified. Because of this lack of

understanding of the factors that influence instability rutting, no effective way of predicting and

evaluating the rutting potential of asphalt mixtures exists today. There is a need to identify the

critical factors that may contribute to instability rutting. Once identified, current laboratory

testing devices can be evaluated or new tests designed to determine a mixture's instability rutting

susceptibility.

Hypothesis

Instability rutting is primarily the result of high near-surface shear stresses at low

confinements. The high near-surface shear stresses at low confinement within the asphalt









pavement driving rutting is the result of the stress distribution in today's high inflation radial

tires. Instability rutting may therefore be controlled by the shear strength and stiffness of the

material at low confinement. The shear strength and stiffness of the materials at low

confinement are influenced by a combination of material state (density), loading history

(compaction), and environmental conditions (temperature). Thus, a laboratory test that can best

replicate this set of critical conditions could be used to determine key material properties for the

evaluation of instability rutting in asphalt mixtures.

Objectives

The main objective of this research is to identify the mechanism of instability rutting at

critical conditions that can lead to a laboratory instability rutting prediction procedure. The only

laboratory test that can replicate high shear stress at low confinement is the hollow cylinder test.

The primary objectives of this research study are listed below:

* Evaluation of instability rut progression under controlled field conditions

* Identification of material stress states under actual tire contact stresses

* Design and construct a new laboratory testing device (hollow cylinder test) to induce
realistic stress states on laboratory asphalt concrete specimens

* Verify laboratory rutting response to field rutting response

* Identification of material response under critical loading and material conditions using the
hollow cylinder device

* Recommend testing configurations and methods for mixture rutting evaluation.

Instability rutting is a complex phenomenon in the area of asphalt pavement distress. The

culmination of these objectives will eventually, after further research and experimentation, lead

to the development of a screening tool or design methodology that can predict the amount of

instability a mixture may undergo under typical loading and environmental conditions. The

research presented here is a first attempt to begin to "get our hands around" the very complex









phenomenon of instability rutting and help pave the way for our ability to predict and understand

instability rutting.

Scope

The research focuses on identifying the critical conditions that contribute to the mechanism

of instability rutting. Defining the conditions that initiate and propagate instability rutting will

lead to the development of better prediction tests and models. It is not possible to examine all

possible parameters that affect rutting of asphalt concrete mixtures in the allotted time frame.

Research Approach

This research is divided into two parts the analytical and the experimental. The analytical

part includes a three-dimensional finite element analysis (FEA) of radial tire contact stresses and

an evaluation of the critical stress states believed to be driving instability rutting. The

experimental part includes the creation of a laboratory testing device, the testing, and data

analysis. A research-approach outline is presented below:

Literature review: examine existing ideas, concepts, theories and results published
on tire contact stresses, rutting in asphalt pavements, torture test devices and
laboratory test devices used in rut prediction

Field evaluation: examine the rut progression of two mixtures under High Vehicle
Simulator (HVS) loading conditions and under two different loading conditions to
understand the propagation of instability rutting under quasi-actual conditions.

Tire contact stresses: employ three-dimensional finite element modeling to estimate
the near-surface stress states in asphalt concrete under radial tire contact stresses.

Laboratory testing device: design and construct a new laboratory-testing device,
the hollow cylinder test (HCT), to induce stresses that would be representative of the
actual stresses induced by radial truck tires from the tire contact stress study.

Test new device: test two mixtures with known field and laboratory performance -
good and poor in the HCT device and evaluate its ability to produce reliable results

Evaluation and observation of tests results

Discussion on the on the results obtained and recommended future testing methods









CHAPTER 2
LITERATURE REVIEW

This chapter reviews some of the literature available on the subjects of permanent

deformation, tire-pavement interface stresses, accelerated pavement testers, torture test devices,

and laboratory methods for predicting mixture performance.

Permanent Deformation

Rutting is one of the major modes of distress in asphalt pavements today. Rutting is

defined as the longitudinal depression that forms within the wheel-path. Rutting can lead to the

premature deterioration of roadways requiring costly rehabilitation and can be a significant

safety concern with skidding and/or hydroplaning when water collects in the ruts.

There are three types of rutting: wear, structural, and instability [Dawley et al. 1990].

The first is wear rutting, which is due to progressive loss of coated aggregate particles from the

pavement surface. This is also known as stripping. The second is structural or consolidation

rutting. This is the traditional term used when discussing rutting. It refers to volumetric

compression and/or shear deformation of the base or subgrade with an assumption that the

asphalt concrete layer contributes very little to the overall rutting of the pavement system --- the

bituminous layer conforms to the shape of the lower layers [Huang 1993]. This mode of rutting

may result from possible insufficient compaction of base and subgrade layers, which undergo air

void reduction and shear deformation under repeated traffic loadings. It can also be due to the

consolidation phenomenon in clayey bases and subgrade. Structural or consolidation rutting will

occur over the design lifetime of the pavement system and is not typically premature failure

mode -- unless the base and subgrade are poorly compacted. Rutted roads due to this mechanism

(Figure2-1) are marked by shallow sloping ruts that are fairly wide (30- 40 inches) [Huber 1999].











original
profile


,~g~t~E~ YI/ ^ b suBgrade
weak subgrade or underlying laver sbgrad
deformation

Figure 2-1. Structural or consolidation rutting. (Reprinted with permission from Huber, G.A,
"Methods To Achieve Rut-Resistance Durable Pavements," Synthesis of Highway
Practice 274, Transportation Research Board, National Research Council,
Washington D.C., 1999.)

The third type of rutting is instability rutting. Instability rutting is due to lateral

displacement of material within the asphalt concrete layer only. Instability rutting is a near-

surface phenomenon occurring in the top 2 inches of the asphalt layer [Dawley et al. 1990].

Instability rutting occurs when the structural properties of the compacted pavement are

inadequate to resist the stresses from frequent repetitions of high axle loadings. The aggregates

rigidly translate and rotate within the asphalt binder [Wang et al. 1999]. Instability rutting

(Figure 2-2) is characterized by steep longitudinal ruts in the pavement with humps of material

on either side of the rut [Huber 1999].




















shear plane

Figure 2-2. Instability rutting. (Reprinted with permission from Huber, G.A, "Methods To
Achieve Rut-Resistance Durable Pavements," Synthesis of Highway Practice 274,
Transportation Research Board, National Research Council, Washington D.C., 1999.)

Tire-Interface Stresses

Over the past thirty years average truck tire inflation pressures have been increasing.

During the AASHO road tests in the 1960s, tire inflation pressures were 75 to 80 psi, today tire

inflation pressures are in excess of 100 psi [Wang et al. 2003]. The type of truck tires on the

roads today is also different form twenty years ago. In the past, trucks utilized bias-ply tires,

which tended to have high wall stiffness and a flexible footprint. Currently, more than 98% of

trucks use radial tires, because of the associated fuel savings and higher reliability of newer tire

structures [Myers et al. 1998, Roque et al. 1998].

Analyses of the tire contact stresses imparted by different tire configurations have been

studied in the laboratory by many researchers. One of the earliest measurement systems

employed electronic pick-ups embedded into the pavement and recorded local forces of a rolling

bias-ply tire [Bode 1962]. In another study, a rotating steel drum was used to measure

automobile tire contact stress [Seitz and Hussman]. It was found in both cases that tire pressures

varied across the tire footprint and longitudinal stresses were present not uniform stress

distribution and only vertical stresses.










Researchers have also measured the normal and tangential tire-pavement interface

stresses by employing a steel-bed transducer array [Woodside et al. 1992]. The Vehicle-Road

Pressure Transducer Array (VSPTA) developed in South Africa measured three-dimensional

tire-pavement interface stresses with 13 triaxial strain gauge steel pins (spaced 17mm

transversely) mounted on a steel plate and fixed flush with the road surface [De Beer et al. 1997].

Some results from VSPTA study (Figures 2-3 and 2-4) display the non-linearity of the vertical

and lateral stress distribution under a radial tire.


-- _- r r - -




. .
0. 5
a



> front A 0 15
Longitudinal rear Lateral


Figure 2-3. Vertical stress distribution of a radial tire. (Reprinted with permission from De Beer
M., C. Fisher, and F. Jooste, "Determination of Pneumatic Tire/Pavement Interface
Contact Stresses Under Moving Loads and Some Effects on Pavements with Thin
Asphalt Surfacing Layers," Proceedings of the Eighth International Conference on
Asphalt Pavements, Seattle, Washington, 1997, pp.179-226.)














I .t .
,- I
* Tllb


a


0 0


Lateral


Longitudinal


Figure 2-4. Lateral stress distribution of a radial tire. (Reprinted with permission from De Beer
M., C. Fisher, and F. Jooste, "Determination of Pneumatic Tire/Pavement Interface
Contact Stresses Under Moving Loads and Some Effects on Pavements with Thin
Asphalt Surfacing Layers," Proceedings of the Eighth International Conference on
Asphalt Pavements, Seattle, Washington, 1997, pp.179-226.)

Dr. Marion Pottinger of Smithers Scientific Services, Inc. developed another device

(Figure 2-5) to measure tire-pavement interface stresses.


Bed

16 Transducers


Rolling


Coaxial Load and
Displacement
Transducer Detail


'y, 5y


(TX, 8x


az, z


Figure 2-5. The Smithers Scientific Inc. tire contact stress measurement device.


mO 0.2
Q,

. 01 -'i
w

(0
* 0


. -01
cJ









By using triaxial load pin transducers inserted onto a flat steel test track, tire-interface

forces and displacements for vertical, longitudinal, and transverse axes were able to be measured

[Pottinger 1992]. The experimental setup used was also capable of determining the rolling tire

footprint shape through the implementation of a rolling steel treadmill device in which the tire

was held in one location, while the bed was moved longitudinally, causing the tire to roll over a

row of 16 transducers. Stresses and displacements were recorded every 0.20 inch longitudinally

(parallel to wheel path) and every 0.15 inch transversely (perpendicular to wheel path) by

varying the transverse position of the sensors. The measurements provided a higher definition of

actual tire-pavement contact stresses than previously obtained. The VRSPTA also measures

contact stresses in the x, y, and z directions, but uses only 13 triaxial strain gauge steel pins,

mounted on a steel plate and fixed flush with the road surface.

Based on contact stress measurements, researchers have identified two distinct types of

contact stress effects that exist under truck tires. These are generally referred to as the pneumatic

effect and Poisson's effect [Pottinger 1992, Myers et al. 1998]. The overriding effect induced

under bias-ply tires is the pneumatic effect, and under radial truck tires is the Poisson's effect.

This is a direct result of tire construction. Radial tires are constructed to have stiff treads and

flexible sidewalls, to minimize the deformation of the tire during rolling. Thus, the lateral

stresses induced on the road by the radial truck tire will tend to push out from the center of the

tire ribs. In contrast, bias-ply tires tend to have high wall stiffness and a flexible tread, resulting

in smaller lateral contact stresses.

Because of the complexities involved in measuring contact stresses under tires, it is not

possible to obtain these measurements directly on real pavements. In particular, the question

arises as to whether stresses measured under a tire on a rigid foundation with embedded sensors









are similar to the contact stresses that a flexible pavement will experience. This question was

answered by a two-dimensional finite element model of a radial tire [Roque et al 2000], using the

finite element program ABAQUS [Hibbit et al. 1997]. The results showed that contact stresses

were nearly identical whether the contact surface was rigid or had properties similar to an asphalt

pavement. Effectively, the stiffness of the tire is so much less than the pavement, that the

resulting contact stresses are similar to those obtained from a rigid steel bed. It was concluded

that contact stress measuring devices with rigid foundations, such as Pottinger's device, are

suitable for the prediction of response of flexible highway pavements.

Thus, radial tires impart not only vertical stress to the surface, but also significant lateral

stresses. This to shift radial tires and their increasing tire pressures and the lateral stresses they

induce is theorized to be a component of the greater prevalence of instability rutting in recent

years [Myers et al. 1999, Drakos et al. 2001].

Accelerated Pavement Testers

Accelerated pavement testing is the application of a controlled prototype load to a

controlled prototype or actual pavement system. Full-scale accelerated pavement testing began

with the AASHO road test in the 1960s. These tests employ actual roadways under full-scale

loadings. They require a lot of space to construct such tracks or available roadways, but can

analyze many different pavement systems and are realistically loaded by vehicles. This method

of testing continues today with the WESTRACK test track and the NCAT test tracks. The

disadvantage of this type of testing is that it also does not allow for quick evaluation of results.

To limit the amount of space needed for full-scale testing and for faster analysis of results

Accelerated Pavement Testers (APTs) were designed.

APTs apply a simulated load to a pavement system. APTs may either be stationary or

mobile devices that apply repeated wheel loadings in a linear or circular manner. Testing under









these accelerated conditions allows for the quick evaluation and response of a pavement system

[Metcalf 1998]. The appeal of APTs lies in their ability to produce realistic wheel loadings,

including wheel wander, to a pavement system. APTs can simulate twenty years or more of

traffic in a reduced period of time.

There are many types of APTs in the world today. These include the Simulated Vehicle

and Loading Emulator (SLAVE), the Texas Mobile Load Simulator (TxMLS), Accelerated Load

Pavement Facility (ALPF) Tester, Accelerated Loading Facility (ALF) Tester, and the Heavy

Vehicle Simulator (HVS). DYNATest, a South African company, developed the HVS. The HVS

consists of a large carriage where a wheel rolls underneath the carriage along the pavement. The

load is applied in a linear manner, and can provide lateral wheel wander and one-way or two-

way directional wheel loadings. The HVS is fully mobile [Kim 2002]. An extensive program was

carried out in South Africa using the HVS. The HVS was brought to different roadway sites in

South Africa to test different pavement sections. The vast amount of data collected through this

testing program formed the basis for the South African pavement design system [National

Institute for Transport and Road Research 1985].

In order to conduct a full-scale test with any APT device, a test track must be constructed

or an available roadway must be available. APTs are also expensive to manufacture (between 1

to 2 million dollars), thus are in limited quantity and require large funding. The primary

disadvantage of APT is this expense and the need to construct appropriate testing facilities.

Furthermore, aging effects are limited, and environmental conditions are not controllable without

special facilities. Despite these limitations, the insight gained with proper measurements and

analysis is worthwhile and can evaluate the effectiveness of other tests.









Torture Test Devices

Torture testing of asphalt specimens has been gaining wide appeal recently for their

relative ease and simplicity. These tests subject an asphalt concrete laboratory -prepared

specimen to repeated loadings. One method of load application is through small wheel loadings

rolling over an asphalt specimen. Torture tests of this method are known as Loaded Wheel

Testers (LWTs). LWTs can be viewed as a small scale APTs. The primary purpose of LWTs is

to perform efficient, effective, and routine laboratory rut proof testing and field production

quality control of asphalt mixtures [Lai 1990]. There are many types of LWTs today. Europeans

have developed the Hamburg Wheel-Tracking Device (HWTD) and the French Pavement

Rutting Tester (FPRT). Americans have developed at the University of Arkansas the Evaluator

of Rutting and Stripping in Asphalt (ERSA) and The Georgia Department of Transportation

(GDOT) began the development of a Georgia Loaded Wheel Tester (GLWT) in 1985. The

Asphalt Pavement Analyzer (APA) is a modification of the GLWT [Lai, 1986]. The concept of

pavement evaluation is simple when using LWTs the rut depth from the LWT can be related to

field performance.

Despite the advantages of simplicity and cost effectiveness, there are some issues.

LWTS can usually, but not always, differentiate between good and bad performing mixtures

[Collins 1996, Stuart and Mogawer 1997]. A comparison of LWTs to APTs showed that LWTS

could distinguish good and bad performing mixtures when only the binder varied. However,

when aggregate gradations were varied, none of the LWTs were able to distinguish between

good and bad performing mixtures [Romero and Stuart 1998]. An analysis of the HWTD and

FPRT noted that the devices could discriminate between mixtures with widely different binder

grades, but failed to give consistent results for mixes with closer binder grades [Stuart and

Mogawer 1997, Stuart and Izzo 1995]. The APA was also sensitive to mixtures with different









asphalt binder and varying gradations [Kandhal and Cooley 1999]. The APA ranked

WESTRACK performance to an 89% accuracy level [Williams and Prowell 1999]. Other issues

of concern include the loading mechanism (pressurized hose or rubber strip or miniature wheel)

the boundary conditions (rigid confinement of the laboratory prepared asphalt concrete specimen

in the testing mold), and the aspect ratio of the loading strip compared to aggregate size [Federal

Highway Administration (FHWA) 1998].

Laboratory Tests

Creep Tests

Uniaxial creep tests are tests where a cylindrical asphalt concrete specimen is subject to a

load in the axial direction with no lateral confinement. This is a single point test -- a single

predetermined loading condition. The load may be applied statically for a predetermined

duration or until a certain deformation has been achieved. The load may also be applied

cyclically in a square wave load or half-sinusoidal loading pattern with rest periods between

loadings. Although traditionally creep means deformation under constant load, the term dynamic

creep is used in pavement mechanics and denotes a load magnitude that is constant but is applied

in intervals.

Creep tests can determine the creep compliance, relaxation modulus, and creep resistance

of asphalt mixtures. Analyses of creep tests provide elastic, plastic, and viscous properties of

mixtures that can be associated to rutting. The uniaxial static creep test has been defended as a

test that is effective in identifying the sensitivity of asphalt concrete mixtures to permanent

deformation with many depth prediction methodologies based on this test [Little et al. 1993,

Marks et al. 1991 and Hills et al. 1974]. The Shell Method is one of the most recognized asphalt

concrete mix design methodologies that was based on the static creep test [Van de Loo 1976].

However, the Shell Method does not incorporate strain hardening with time [Ali and Tayabji









1998] and may overpredict rutting in modified asphalt mixtures [Monismith and Finn 1987]. The

disadvantages of the static creep test are its simplicity and defined loading condition on the

specimen.

In order to better simulate loading conditions found in the field, testing under repeated

load or cyclic conditions was introduced. Many researchers have found good correlations

between repeated creep load tests and field performances and have developed rut prediction

models [El Hussein and Yue 1994, Qi and Witczak 1998]. The creep test is shown to be sensitive

to mixture variables including asphalt grade, binder content, aggregate type, air void content,

temperature of testing, testing stress level, and rest periods [Little et al. 1993]. Any correlations

made with creep testing are very dependent upon the load and rest times [Qi and Witczak 1998].

Comparisons of the asphalt mix stiffness obtained from compressive creep tests with asphalt mix

stiffness obtained from wheel-tracking tests observed good agreement between the two

techniques at high values of bitumen stiffness. However, at low values of bitumen stiffness, the

results from the creep tests gave higher asphalt mix stiffness than those from the wheel tracking

tests [Hills et al. 1974].

Furthermore, creep testing does not allow for a unique separation of plastic strains when

working within the elasto-plasto-visco material framework. Another shortcoming of the uniaxial

compression tests is that no direct separation can be made in response of the material to

hydrostatic and deviatoric stress [Kim et al.]. The major limitation for uniaxial creep testing,

both static and cyclic, is the lack of lateral confinement. Lateral confinement has been seen to be

critical for proper analysis of rutting susceptibility of open graded or mastic mixtures [Little et al.

1993]. Although simplicity is the advantage of this test, the predetermined method of loading,









lack of confinement, and divergent results make this test difficult to employ as a simple reliable

quantitative test for instability rutting prediction.

Complex Modulus Testing

The dynamic or complex modulus test is another test often used to predict the rutting

susceptibility of hot mix asphalt mixtures. Complex modulus testing was first described as a test

on hot mix asphalt in 1962 [Papazian 1962]. Recently, NCHRP Project 9-19 evaluated the

complex modulus test and the AASHTO 2002 Design Project focused complex modulus testing

as a Simple Performance Test (SPT) for the rutting resistance of HMA mixtures [Pellian and

Witczak 2002]. The dynamic modulus test is outlined in ASTM D3497.

In complex modulus testing, sinusoidal stress or strain amplitudes are applied axially to

an unconfined cylindrical specimen at 16, 4, and 1 HZ. The ASTM Standard also recommends

testing at temperatures of 5, 25, and 400C. The complex modulus test is similar to the unconfined

creep testing in testing set-up, but the load application and analysis of the response differs.

The dynamic or complex modulus test relates the cyclic strain to the cyclic stress in a

sinusoidal load test. The complex modulus is defined as:


E*= o
go

Where: oo is the stress amplitude,

So is the strain amplitude.

The complex modulus, E*, is composed of a real component known as the storage

modulus, E', and an imaginary component known as the loss modulus, E". The storage modulus

represents the elastic portion of the response and loss modulus represents the viscous portion of

the response. The storage and the loss modulus can be obtained by measuring the lag in the









response between the applied stress and the measured strain. This lag in the response is known as

the phase angle (6).

When conducting an axial complex modulus test lateral strains can be also be measured.

With lateral and axial strains measured, the Poisson's ratio can be determined. The shear

complex modulus, G*, can be determined by the following equation [Harvey et al. 2001]:

E*
G'=*-
S(1+ 2v)

This equation assumes that Poisson's ratio is constant, although Poisson's ratio has been

seen as being frequency dependent [Sousa and Monismith 1987] and that linear elastic relations

with moduli hold for visco-elastic complex moduli. Recent findings using torsional complex

modulus testing suggest that reasonable values of Poisson's ratio can be determined from E* and

G* [Pham 2003]. Torsional complex modulus testing provides G* directly. Torsional complex

modulus applies a torsional sinusoidal stress or strain amplitudes are applied axially to an

unconfined cylindrical specimen rather than an axial stress or strain amplitude.

The axial complex modulus has been shown to have a good correlation with the rutting

resistance of HMA paving mixtures [Witzak et al. 2002]. Numerous research groups have shown

that the complex modulus test can be used to characterize the temperature dependency of

mixtures and viscosity characteristics over time [Witczak et al. 2002, Perraton et al. 2001].

Analysis methods for the characterization of rutting resistance in asphalt mixtures with the

complex modulus test have been presented [Majidzadeh et al. 1979; Shenoy and Romero 2002].

The complex modulus test is still only a measure of the visco-elastic properties of an asphalt

mixture. Since the interpretation of the complex modulus is based on the assumption of linear

viscoelasticity of the mixture, it is necessary to maintain a fairly low strain level during testing to

avoid any non-linear effects. Maintaining a stress level that result in a strain response that is









close to linear is critical to achieve a test that is reproducible and allow for proper analysis. Strain

amplitudes of 75 to 200 microstrain are suggested in order to maintain linearity during triaxial

compression testing [Witczak et al. 2000]. As a drawback, complex modulus testing does not

look at large strain creep or plastic strains or particle movement that truly capture rutting in the

field.

Triaxial Testing

Triaxial testing allows for many possible stress states on an asphalt specimen. Virtually

any combination of hydrostatic and deviatoric stresses can be induced on a specimen in the

triaxial test [Kim 1997]. Confinement can be varied and depending on the method of loading

tests may either be in compression or extension. Loading can either be strain or time dependent,

as well as static or cyclic.

Early researchers found many similarities between soil and asphalt mixtures when testing

under static triaxial conditions. Both soil and asphalt concrete exhibited cohesion and pressure

dependency [Goetz and Schaub 1951, Nijboer 1948, Gandhi and Gallaway 1967]. These early

triaixal tests were conducted by applying a constant confinement to a cylindrical asphalt

specimen and then loading in the axial direction at prescribed rate. Mix design methods using the

static triaxial test and modeling asphalt concrete as a Mohr-Coulomb material soon followed

[Smith 1949 and McLeod 1952]. Today triaxial testing has been used to determine rutting

susceptibility of mixtures [Krutz and Sebally] by comparing the failure envelopes of different

mixtures and new mixture design techniques still continue using static triaxial testing [Mahboud

and Little 1988]. These tests are limited to measuring modulus and strength under static

compression. These tests are static tests, and are not representative of the repetitive loading

experienced in the field by the asphalt concrete layer. A desire for more realistic loading led to

cyclic or repeated-load triaxial testing.









Repeated-load triaxial tests apply an axial stress that varies with time in manner more

representative of field loading conditions. Various pulse loadings have been used depending on

the available equipment, but generally a sinusoidal or triangular loading pattern is employed for

its close approximation to actual field loading and to avoid sudden stress changes [Brown 1976].

In fact it has been shown that for the same degree of stress, a square loading produces greater

strain than sine wave [Brown and Cooper, AAPT]. This is likely due to the sudden impact of a

square wave compared to a sine wave's gradual build-up.

In the United States, most loading is 20 to 30 repetitions/minute, with a pulse of 0.1

seconds in duration [Brown 1976]. Investigations conducted at the University of Nottingham

suggest that rest periods are not significant for permanent strain tests on asphalt mixtures.

Repeated triaxial testing allows the visco-elastic properties of an asphalt material to be obtained

[Rowe et al. 1995]. These parameters are directly related to rutting and enable predictions of

deformation in the field and comparisons to be made between different materials.

The repeated triaxial testing has been considered by some to be the best practical method

for the testing of pavement materials against rutting [Brown 1976]. This is because of its ability

for various stress and loading conditions that best simulate actual field conditions. Some models

have been proposed for rutting using the repeated triaxial test [Meyer et al. 1976]. Despite the

advantages, little repeated triaxial testing is performed to asphalt concrete [Brown 1976].

Repeated triaxial testing is primarily used for bases and sub-grades. The resilient modulus (MR)

of the subgrade is the basis of the AASHTO asphalt concrete mix design methodology. The

major disadvantage of triaxial testing is that only principal stresses can be applied to a specimen

-- the orientation of principal stresses is always in the vertical or horizontal direction.









Furthermore, the application of high shear at low confinement is not possible the state of stress

hypothesized to be critical to instability rutting.

Superpave Shear Tester

The Superpave Shear Tester (SST) is a closed-loop feedback, servo-hydraulic system that

can apply axial loads, shear loads, and confinement pressures to asphalt concrete specimens at

controlled temperatures. The response of asphalt concrete to these loads can be used as inputs to

performance prediction models [Shenoy et al. 2001]. The SST is a product of the Strategic

Highway Research Project (SHRP) to evaluate asphalt mixture performance. The test was

originally developed at the University of California at Berkeley [Sousa et al. 1991]. The SST was

developed to address the following issues not accounted for in other tests:

* Dilation under shear loading

* Increase in stiffness with increase in hydrostatic pressure,

* Negligible volumetric creep,

* Residual permanent deformation on removal of load,

* Temperature and rate loading dependence, and

* Difference in response between creep and repeated loading.

The SST has six main components: testing chamber, test control system, environmental

system, hydraulic system, air pressurization system, and measurement transducers. The testing

chamber includes a reaction frame and a shear table. The reaction frame is extremely rigid

providing for accurate measurements. The shear table holds a specimen during testing and is

capable of applying shear loads. The specimens normally have a diameter of 150 mm and a

height of 50 mm; however, specimens with diameters and heights up 200 mm can be tested with

only minor modifications to the system. Shear and axial loads can be applied sinusoidally,









repetitively or in static creep mode under a confining pressure and at temperatures anywhere

from 0 to 700C. Four Superpave tests are capable of being performed with the SST:

* Repeated shear at constant height (RSCH)

* Repeated shear at constant stress ratio

* Simple shear at constant height (SSH)

* Frequency sweep at constant height (FSCH)

The AASHTO Provisional Standard TP7-94 contains a detailed description of these tests.


Figure 2-6. SST test chamber.

On-going research is attempting to evaluate the reliability of the SST in rut prediction.

Results of a study conducted in North Carolina indicated that the RSCH test can clearly identify

the well- performing versus poorly performing mixes. The RSCH correctly predicted the non-

susceptibility to rutting of one mix, where the French Pavement Rutting Tester and the Georgia









loaded-wheel tester indicated early rutting of the pavement [Tayebali et al. 1999]. A

comparative evaluation of rutting and stripping of two asphalt mixtures using the SST and the

Hamburg wheel-tracking device (HWTD) has also been conducted. The results showed that both

the SST and HWTD correctly predicted the superiority of the Superpave mixture to the Marshall

mixture [Wang et al. 2001.]. A comparative study of the RSCH to the APA was also conducted.

Correlations of various parameters (permanent deformation or strain, slopes, and intercepts from

linear or power law regressions) from the RSCH had significant correlation with APA rut tests

[Zhang, 2002].

However other research has indicated that the SST and the performance models

developed with this test have some errors that made prediction impossible [Anderson, et al.,

1999]. Analysis of the RSCH test indicated that the coefficient of variation of the permanent

strain at 5, 000 load cycles for samples was between 10% and 20% with a suggestion that the

RSCH be used for low volume road prediction only [Romero and Anderson 2001]. Other

researchers have also indicated high variability when testing with the SST and results depended

on the analysis method [Romero et al. 1998]. Further research has found that the SST could

differentiate between asphalt binders but are not sensitive to changes in aggregate [FHWA

1998].

Hollow Cylinder Testing

The hollow cylinder test (HCT) has been used since the 1930s to study soil specimens

subject to pure shear. Since then, the HCT has become more versatile with researchers now

applying axial, torsional, and internal and external pressures [Saada and Townsend 1981]. The

HCT has been used most extensively in soil testing. The versatility of stress states capable with

the hollow cylinder has made it ideal for the studying of the anisotropic nature of soil. The

rotation of principal stresses and torsional properties of materials has been shown to have









significant effects on the strength and stiffness of soil [High et al. 1983, Miura et al. 1986]. There

is a plethora of articles and literature in geotechnical engineering on hollow cylinder testing,

configuration, issues, and material properties [Hight et al. 1983, Alarcon et al. 1986, Dusseault

1981, Vaid et al. 1990, Richardson et al. 1996]. Only recently, American pavement researchers

have begun to utilize the hollow cylinder for the analysis of asphalt concrete for the

determination of tensile, creep, and dynamic properties [Sousa 1986, Crockford 1993, Buttlar et

al. 1998, Alavi and Monismith 1994].

The University of California at Berkeley conducted hollow cylinder testing on asphalt

concrete using specimens with an inner radius of 3.5 inches (88.9 mm) and an outer radius of 4.5

inches (114.3 mm). The specimen height was originally 18 inches (457.2 mm). Specimens were

created through the use of a special fabricated molds, compaction device, and unique methods

[Sousa 1986]. These same researchers decreased the height to 8 inches (203.2 mm) with

specimens compacted through a rolling wheel compaction procedure outlined in SHRP A-003.

Research was done using this device to investigate the time and temperature dependent

properties of asphalt concrete mixes and to examine the applicability of linear viscoelasticity to

asphalt concrete material response. Specimens were subject to axial and shear frequency sweeps

from 10 Hz to 0.01 Hz and at temperatures of 4, 25, and 400C. During the axial sweeps only an

axial stress was applied and during the shear sweeps, beside a small seating load, only shear

stress was applied. Stress levels were selected in this study to be as small as possible to allow for

nondestructive testing and for the material to remain in the elastic range [Alavia and Monismith

1994].

The United Sates Air Force also conducted research using the hollow cylinder device.

This investigation was driven by rutting occurring in the asphalt concrete runways of F-15 jets. A









hollow cylinder specimen was used in the laboratory study of asphalt concrete subjected to

realistic (but very slow moving) vehicle load simulation. A statistically significant difference is

illustrated between the permanent strain resulting from laboratory testing with no principal-plane

rotation and that resulting from testing with principal plane rotation. The tests conducted with the

hollow cylinder were able to identify statistically significant differences between two levels of

wheel loading, two aggregates, and two asphalts [Crockford 1993].

Another researcher today is using the hollow cylinder to obtain fundamental properties of

asphalt mixtures such as creep compliance and tensile strength at low and intermediate

temperatures. Hollow specimens used in this research have a six-inch (150mm) outer diameter

and a 4-inch (100mm) inner diameter, giving a wall thickness of one inch (25 mm). Hollow

specimens were created by drilling a hole through a gyratory compacted solid specimen. In this

test, pressure is applied to the inner cavity resulting in tangential tensile stress or "hoop" stress

developing in the wall. This test hopes to become a surrogate test for the Superpave Indirect

Tensile Test (IDT) in asphalt mixture design [Buttlar et al. 1998].

The stress states that can occur in the hollow cylinder are almost unlimited when one has

the ability to vary the internal and external chamber, the axial, and torsional stresses. The HCT

device allows for the rotation of principal stress and planes not allowed in other laboratory tests,

and can apply internal and external pressures unlike the SST.

The following equations help to illustrate the variation in stresses across a hollow

specimen when subject to stresses in three directions. These stresses are based on the theory of

linear elastic thick-walled cylinders. All the notations in the equations are based on Figures 2-7

and 2-8, with stresses are assumed to be uniform along the entire length of the sample.






















Figure 2-7. Plan view of a hollow cylinder with outer radius, ro, and inner radius, ri, and subject
to varying outer (Po) and inner pressure (Pi).




)Torque at
Rotation







Y / r nax
HMA
Specimen

S/ Y= Single
Shearing








Rigidly Fixed
at Bottom


Figure 2-8. Profile view of a hollow cylinder specimen.









The following equations are for the radial (or) and tangential (Gt) at any radius (r) across

the specimen subject to different internal (pi) and external pressures (po) based on the outer (ro)

and inner (ri) radii [Singer and Pytel 1980]:

rP rp-rPo r^ ro(p -po)
"r o 2 2 2 2
r- r, r, -r, P
i _r2p,-rpo ro0(p1-po)
t 2 2 2 2 2
r2 -r,2 r o -r,

The average radial and tangential stresses across the specimen can be calculated and

reduced to the following equations:

((p r + p r)
r, (average) = p0r P
(r0 +
(po Pr)
a, (average) = (oro 0 Pr)
(ro -r)

The location of the average stress is based solely on geometry and occurs at:

r = rr
average o

The shear stress (c) varies linearly across the specimen. The following equation is for the

shear stress at any radius (r) in the specimen [Ni 1987]:


r(r)= > max X
r
Tx r
max J


2

The average shear stress across the specimen is given as:









T
r =r -
average eq

2 (r~-r
r =-X 0-
eq 3 r2 2)

The vertical stress is assumed to act uniformly over the specimen; the equation for the

vertical stress is given below [Hight et al. 1983]:

W (par2 pr22
vertval 2 2 2 2
r2r -r r
0 I 0 I
-O -


As shown above stress states across the specimen are not uniform. A hollow cylinder

specimen should be thin enough to allow a uniform of stress distribution, but thick enough to

contain a representative sample and provide for the maximum particle size. The length should be

adequate to minimize end effects. Most guidelines established for hollow cylinder testing have

been based on soil testing. For instance, as a rule of thumb, it has been suggested to have a wall

thickness to grain size ratio of 10 to 25 when testing in the hollow cylinder [Hight et al. 1983].

This is to ensure a representative sample. This is not a tremendously difficult when testing clays

or fine sands but unreasonable for testing asphalt concrete where aggregate sizes may approach

one inch. Past researchers have used a one-inch wall thickness when testing asphalt concrete

with maximum aggregate sizes of 38" to 3/4" (9.5 to 19 mm) [Buttlar et al. 1998 Alavi and

Monismith 1994].

Guidelines have been established for the length and wall thickness based on equations

from the theory of thin elastic cylinders [Saada and Townsend 1981]. (A thin walled cylinder is

one where the stress distribution across the wall can be considered uniform). These equations

were established to ensure that normal stress non-uniformity across the sample is kept to a

minimum. However, these equations were based on the assumption of a thin cylinder. This









assumption is reasonable for soil testing where the hollow cylinders could be considered thin. If

a hollow cylinder is assumed to be thick, these equations break down and actually suggest that a

very thick specimen is just as desirable as a very thin specimen. A finite element analysis is

typically employed [Hight et al. 1983, Vaid et al. 1990, Buttlar et al. 1998] to quantify stress

concentrations and non-uniformities for different geometries and provide insight into end effects.

The finite element analysis provides boundary conditions not possible through the theoretical

equations that assumed free boundaries and a thin cylinder.

The issue of shear stress non-uniformity across the specimen does not depend on weather

one considers a thin or thick specimen since the shear stress varies linearly across the specimen.

The maximum shear stress occurs on the outside edge and the minimum shear stress occurs on

the inside edge. The level of shear stress non-uniformity across a specimen is typically quantified

with the following non-uniformity coefficient [Timoshinko and Woinowsky-Krieger 1959]:

1 1 ro
3 = ag dr
ro r, Tavg r,

One researcher has suggested a 33 value of less than 0.11 is acceptable [Vaid et al. 1990].

The important issue, as with any laboratory test, is to consider the actual stress state on the

specimen by understanding the geometric configuration of the specimen and the boundary

conditions.

Summary

The discussion presented indicates there are many laboratory tests today that are currently

being used to evaluate a mixture's performance in the field. Some are large-scale and others

small-scale. Some induce many stress states on the asphalt concrete laboratory specimen and

others attempt to induce a single stress state on the asphalt laboratory specimen. The hollow









cylinder device is a test that can achieve multiple stress states and stress paths as well as

principal plane orientation not capable in other laboratory tests.









CHAPTER 3
HEAVY VEHICLE SIMULATOR TESTING ON ASPHALT PAVEMENTS

The opportunity to analyze the progression of instability rutting under controlled field

conditions became a reality when the Florida Department of Transportation (FDOT) along with

the South African Council of Scientific and Industrial Research (CSIR) purchased a Heavy

Vehicle Simulator (HVS) Mark IV [Kim 2002]. Dynatest, a South African company,

manufactures the Mark IV.




























Figure 3-1. The Mark IV Heavy Vehicle Simulator device.

The Accelerate Pavement Testing (APT) facility was constructed in Gainesville, Florida, to

test the performance of asphalt mixtures using the HVS. The insight gained by this testing and

analysis was valuable in understanding the phenomenon of instability rutting.









Instability Rutting in Modified and Unmodified Pavements Under HVS Loading

A testing program to evaluate the performance of modified and unmodified asphalt mixes

under HVS conditions was instituted by the FDOT. Modified asphalt mixes refer to asphalt

mixes where the binder used contains a polymer modifier. The polymer-modifying agent

employed in this study was styrene-butadiene-styrene (SBS). The HVS employs a super-single

radial tire with an average contact stress of 115 psi and a footprint of 12 inches wide by 8 inches

long. The load is applied uni-directionally at a speed of 6 mph. The testing was performed at a

uniform pavement temperature of 500 C made possible by an environmental control chamber.

The modified and unmodified asphalt mixtures used in the study were both fine graded SP-12.5

mixtures, with the unmodified binder rated 67-22 and the SBS-modified binder rated 76-22

(Superpave nomenclature).

The pavement system at the APT facility consisted of a 4-inch layer of asphalt concrete, a

10.5-inch limerock base, a 12-inch stabilized limerock subgrade, and a natural sandy subgrade

(A-3 soil). The asphalt concrete layer was placed in 2 two-inch lifts with initial air void contents

of 7%. A detailed description of the testing program and results has been published [Tia et al.

2002, Tia et al. 2003, Byron et al. 2003].

Rutting Depths

Rut depths were measured by a laser profiler. The laser profiler consists of 3,538 lasers

attached to a movable frame. Each laser is accurate to within 0.1 mm. The profiler has 59 rows

of lasers in the longitudinal direction (parallel to the direction of travel) spaced every 4 inches

with each row consisting of 60 lasers spaced every inch in the transverse direction (perpendicular

to direction of travel). Laser profile measurements were taken incrementally after a certain

number of passes from the HVS. A transverse profile of movement of the pavement surface, i.e.

rutting, was then constructed. A straight-line was drawn over the surface profile tangent to the









highest points. The greatest distance from the straight-line to the trough was taken as the rut

depth. This method is known as the surface profile method and is the method most used in

defining rut depth [Kim 2003]. Measurements were generally taken after every 100 passes up to

1,000 total passes, then every 500 up to 3,000 total passes, then every 1,000 to 5,000 until

100,000 passes or the rut depth was deemed sufficient.

The results of the modified versus unmodified mixes at 500C demonstrated that modified

mixtures do a better job at resisting instability rutting than unmodified mixtures. Five sections

were tested and the descriptions of the each section are described below.

Table 3-1. Test sections in HVS testing.
Test Temperature
Section Mixture Type Symbol in Figures
(oC)

1B 50 Modified U

2B 50 Modified U

4A 50 Unmodified 0

4B 50 Unmodified 0

5A 50 Unmodified 0



Sections 4A, 4B, and 5A (unmodified mixture sections) had about two and half times the

rut rate as compared with that of sections 1B and 2B (modified mixture sections).











20

18 4B
4A
16-





10
5A







6

4
2
tio ~~ ,y-------------^ --








0 50000 100000 150000 200000 250000 300000
Number of Passes
Figure 3-2. Rut depth progression for the various mixes.

The results indicate that both the modified and unmodified sections have an early response

(fist 1,000 passes) that is very similar. After about 1,000 passes, the sections begin to display

differences. The modified sections' rut rate decreases and achieves a stable linear progression.

The unmodified sections' rut rate continues at the same rate and does not reach a stable rut rate

seen in the modified sections.

Trench cuts taken after testing (Figures 3-3 and 3-4) had been completed displayed no base

or subgrade deformation, indicating true instability rutting.

































Figure 3-3. Trench cut on one of the sections tested during the HVS study.


Figure 3-4. Transverse cross-section of the asphalt pavement after testing.













Rut Propagation Analysis through Surface Profiling

The laser profiler provided detailed measurements of the surface movement of the

pavement section. Figure 3-5 displays an average transverse profile of one of the sections after

100 and 1,000 passes and Figure 3-6 displays the range of the longitudinal measurements that

were averaged to get a single point in Figure 3-5.



2.0 1,000 Passes

-e-100 Passes

1.0 4 -



0.0
E
S-1.0
I-
0




-3.0 Transverse
Location
Investigated
-4.0
-30 -20 -10 0 10 20 30
Average Transverse Cross-Section Profile (inches)

Figure 3-5. Average transverse profile created by averaging the 58 longitudinal measurements
from the laser profiler.











0.0

-0.5

-1.0


E -1.5
0
-2.0

-2.5

-3.0
Average Value
-3.5 Reported for
Transverse Profile

-4.0
0 40 80 120 160 200 240
Longitudinal Cross-Section (inches)

Figure 3-6. The 59 longitudinal values averaged.

The apparent variability in the longitudinal profile was not significant since the scale is in

millimeters. The slight bulge in the middle of the wheel-path in Figure 3-5 is not surprising and

is often seen in granular material under footings. The bulge occurs since the material in the

center of the loaded area is under confined vertical compression and cannot deform laterally,

thus the vertical movement is limited primarily to one-dimensional vertical movement.

With the high degree of definition of the surface deformation from the laser profiler, a

careful analysis of the progression of the rutted surface was initiated. From each average

transverse profile, a ratio was calculated of the total surface area elevated over the total surface

area depressed after each measurement. This ratio would help to determine the differences in

instability rut progression and response between modified and unmodified mixtures. These

measurements are only based on surface changes in height and depth and cannot reflect the










changes in densities or air voids. Figure 3-7 displays the method used in calculating the surface

area ratio from each transverse rut measurement.


2.0


1.0


( 0.0
E
0
E Area 3

*-1.0 Area 1
1- --Average Transverse
4 Profile after 1,000
S-2.0 Passes -


-3.0
-Ratio Used in Presentation:
[(Area 2+ Area 3)/Area 1]*100%
-4.0
-30 -20 -10 0 10 20 30
Average Transverse Cross-Section (inches)

Figure 3-7. Average transverse profile with ratio equation.

Figure 3-7 displays the deformation of the depressed (Area 1), elevated (Areas 2 and 3),

and unaffected outside areas of the transverse cross-section from the laser profiler measurements.

Figure 3-8 displays the ratio of the surface area elevated to surface area depressed for five

sections presented above. The ratio was calculated for every rut measurement -- cumulative

rather than incrementally. (Although rut rate could have been analyzed as well in a similar

manner, assurance that the laser profiler was in the exact location for each measurement coupled

with very small incremental changes made the degree of error greater than the degree in accuracy

and thus was not attempted.) The logarithmic scale was selected only because it leads to a

straight-line plot.











100
90

80
70
(- 4A 5A




10
0
) 40 | I -
< -^ ---* 2B






100 1000 10000 100000 1000000
Pass #

Figure 3-8. The ratio of area elevated to area depressed for all 500 C sections.

The ratio of elevated surface area to depressed surface areas is roughly the same for both

the modified and unmodified sections in the first 100 passes, but the unmodified sections diverge

from the modified sections at 200 passes for Sections 4A and 4B and at 500 passes for Section

5A. The ratio of the area elevated to area depressed remains constant throughout testing for the

modified mixtures at around 40%. The ratio of the surface area elevated to surface area

depressed for the unmodified mixtures increases rapidly as the testing proceeds, reaching almost

90% for Sections 4A and 4B at 100,000 passes. These results may suggest that rutting in the

modified sections was primarily due to vertical densification where the material within the wheel

path is primarily compressed with some material laterally displaced. The rutting in the

unmodified sections was due to vertical densification but also significant lateral

shoving/displacement. This hypothesis that the instability rutting in the modified sections was

primarily vertical densification and the unmodified sections was primarily lateral

shoving/displacement was investigated further through the analysis of asphalt cores taken after

the HVS testing.











Volumetric Analysis of Asphalt Cores

The wheel-path was 16 inches wide 12-inch tire width plus 4 inch tire wander. Cores


were extracted from the wheel-path (+5 inches from the travel centerline), the elevated region


(14 to 16 inches from the travel centerline), and from 24 inches outside the travel centerline -- an


area deemed undisturbed by the rutting -- from Sections 1B and 5A. Figure 3-9 displays the final


transverse rut profile of Sections 1B and 5A after 110,000 and 100,000 passes, respectively.




5A
6

4

E2 2
E 0

o0


0







-30 -25 -20 -15 -10 -5 0 5 10 15 20 25 30
E l

04 1
4-e
-6

-8

-10
-30 -25 -20 -15 -10 -5 0 5 10 15 20 25 30
Average Transverse Rut Profile (inches)

Figure 3-9. Final transverse rut profile of sections 1B and 5A.

The each core was divided into Lift 1 (first or bottom asphalt lift) and Lift 2 (top asphalt

lift) by the FDOT. The following table summarizes the air void content and heights of the cores.










Table 3-2. Summary of asphalt cores from pavement sections 5A and 1B post-testing.
Height per Region Air Voids per Region
Section Layer (mm) (%)
Wheelpath Elevated Outside Wheelpath Elevated Outside
Lift 2
ft2 41.3 46.2 45.3 4.46 8.39 6.75
5A Lift 1
5A Lft 43.9 51.5 49.7 2.64 4.88 4.25
Total/Average 85.1 97.7 95.0 3.52 6.54 5.44
Lift 2
37.9 40.7 39.6 5.14 7.02 6.67
1B Lift 1
B Lft 48.2 51.9 51.1 3.77 4.70 6.07
Total/Average
Total/Average 86.1 92.6 90.7 4.37 5.72 6.33


The change in total height from the wheelpath to the outside and from the elevated region

to the outside from Section 5A is negative -9.9 mm (downward movement) and +2.7 mm

(upward movement), respectively. These values can be seen in Figure 3-8 and shows that the

wheelpath core was likely taken 4 inches to the Right of the Centerline and the elevated core was

likely taken 14 inches to the Right of the Centerline.

The change in total height from the wheelpath to the outside and from the elevated

region to the outside from Section 1B is negative -4.6 mm and +1.9 mm, respectively. These

values can be seen in Figure 3-8 and shows that the wheelpath core appears to have been taken

from approximately 4 inches Right of the Centerline and the elevated cores appear to have been

taken from approximately 14 inches Right of the Centerline the approximate same offsets as

the cores from Section 5A.

The approach taken to clearly and concisely illustrate the primary mode of instability

rutting each section underwent was through phase diagrams based on the following concepts.









Each core is assumed to be composed of two volumetric phases a solid phase and void

(air) phase. The solid phase consists of the aggregate and binder and the void volumetric phase

consists of air. Air void content (AV%) can therefore be defined below as

Vv
AV%=-- x 10
V

where V is the total volume, or the sum of the Vv (air) and Vsolid and .

For each core, the volumetric phases can be simplified into heights the cross-section
area is the constant of all the cores. Therefore, the total height (H) of each lift consists a certain
height of air (Ha) and solids (Hs), or as follows:
H = Ha+ Hs

And therefore,

AV% = Ha/H.


Figure 3-10. Phase diagram of the asphalt cores.

With this simple concept and the air voids and heights of the cores, an analysis of the
cores can be undertaken clearly and simply. The Table 3-3 provides the heights and change in
heights for each phase of each of the lifts of Sections 1B and 5A.











Table 3-3. Summary of core data through use of phase diagrams.
Initial Wheelpath Phase Elevated Phase Change
Section Lift Phase Heights Region Change of Region of Total
(mm) Heights Total Height Heights Height
(mm) Change (mm) Change
Air
3.06 1.84 30% 3.87 96%
Two Solid
42.26 39.42 70% 42.29 4%
5A Air
2.11 1.16 16% 2.52 22%
One Solid
47.61 42.72 84% 49.03 78%
Air
2.64 1.95 40% 2.86 20%
Two Solid
36.98 35.94 60% 37.83 80%
lB Air
3.10 1.81 44% 2.44 -79%
One Solid
47.98 46.37 56% 49.48 179%


The percentages reveal the story. If the entire change in height change in the wheelpath

was due to vertical densification, then the percent loss attributed to air would be 100%; and vice-

versa if the entire change in height was due to shoving then the percent loss attributed to solid

loss would be 100%. (The percent loss of air and solid must add to 100%.)

The wheelpath cores for Section 5A indicate that only 16% and 30% of the height

change was due to air void loss densificationn) with 70% to 84% due to solid material loss

(shoving) in Lifts 1 and 2, respectively. However, the cores from Section 1B indicate an almost

an even 50/50 split in the loss of height due densification and shoving.

The percentages from the elevated region present a much more interesting phenomenon.

Section 5A in Lift 2 indicates very little gain in solid material (4%), but a significant gain in air

(96%). Section 5A Lift 1 the gain in height is primarily due to solid material (78%). Section 1B









Lift 2 is very similar to Section 5A Lift 1 with 78% of the height gain coming from solid

material. However, Lift 2 actually loses air voids with the infusion of solid material (-79%).

This suggests that the elevated region in both sections gain in height by material being

shoved away from the wheelpath. Once the material is displaced laterally from Section 5A

(unmodified section) the material appears to expand or possibly dilate, while in Section 1B

(modified section) the expansion or dilation of this material is not a significant contributor to

elevation in height rather.

One-way Versus Two-way Directional Loading

The FDOT also conducted an evaluation on the most efficient method of HVS loading.

This study investigated one-way versus two-way directional tire loading. The results clearly

showed that one-way directional loading resulted in faster rut progression and in greater rut

depths. The uni-directional loading caused the rut to develop at a rate of approximately 65

percent greater than that of the bi-directional loading. The results of this study are presented

below.














I 7C(Bi-directional No wander) 7B-W(Uni-directional No wander)
16

14

12

E 10






4

2



0 100000 150000 0200000 250000 300000 350000
Num ber of Passes


Figure 3-11. Comparison of one-way and two-way directional loading with no tire wander.



7BE(Uni-Directional with 4-inch Wander) 7AEc Bi-Directional with 4-inch Wander





7





4



2
I



0 100000 200000 300000 400000 500000 600000 700000 800)00 900000

Number of Passes

Figure 3-12. Comparison of one-way and two-way directional loading with tire wander.


It is important to note that the one-way loading sections had slightly lower average


temperatures than the two-way loading sections [Tia et al. 2001, Tia et al. 2002, Byron et al.


2003]. Differences in rutting from one-way or two-way directional loading have not been









observed at low temperatures [Huhtala, M. and J. Pihilajamaki 2000]. It is important to note that

the rutting that occurred was indeed instability rutting. Trench cuts conducted after the one-way

and two-way directional loading showed that there had been no deformation of the base layer;

rutting was confined to the asphalt layer only i.e., true instability rutting similar to Figure 3-3

and 3-4. Figure 3-13 below illustrates directional axes inserted for illustrative purposes and to

define terms used below.


Figure 3-13. Instability rutting with directional axes.

The directional testing results raised two questions. First, it is well known that asphalt

concrete exhibits visco-elastic properties [Ullidtz 1987]. In two-way directional loading, the

relaxation time is about half that of the one-way directional loading (tire load must be picked up

and moved back to the starting position after each pass in one-way directional loading). A greater








relaxation time should provide time to recover more strains, but actually the one-way directional

loading with greater relaxation times between passes produces the greater rutting. (Not to

mention that the one-way directional loading sections had an average temperature lower than the

two-way directional loading sections). The following figure illustrates the additional time that is

required during one-way directional loading compared to two-way directional loading.







O-0 vs. 0

Two-Way Loading One-Way Loading
Figure 3-14. Two-way versus one-way directional loading.
Second, if instability rutting is thought to be due to shoving near the surface caused

primarily by transverse shear stresses (Plane YZ) near the edges of the wheel-path (Huang 1993,

Drakos et al. 2001, and Huber 1999], then there should be little difference in the rut depths

between one and two-way directional loading. Since the stress paths in the YZ-Plane will not be

different from one-way and two-way directional loading. This is described in Table 3.4 (the

letters A, B, and C, will be used for illustrative purposes).











Table 3-4. Deformation of an element in an asphalt pavement near the edge of the tire in the YZ-
plane (transverse plane).
Shape of Element in AC Shape of Element in AC
Layer, Left Side of Tire Point of Interest Layer, Right Side of Tire
Edge Edge
Far Away, Approaching
(A)
Next to,

(B)
Far Away, Passed
(C)



The above table illustrates that shear stresses that cause the deformation will not discern

between one-way and two-way directional loading. Two passes of the wheel load in one-way

loading will produce a shear deformation pattern of A-B-C A-B-C; two passes of the wheel

load in two-way loading will produce a shear deformation pattern of A-B-C C-B-A. The

magnitude of B would not vary between passes or upon the one-way or two-way directional

loading.

Stresses in the longitudinal plane (plane XZ) must also have an affect on the degree of

rutting since these stress paths along this plane will vary depending on whether the load

application one-way or two-way directional loading. The deformation of an element in the

asphalt pavement near the tire edge in the XZ-Plane (longitudinal Plane) is illustrated in Figure

3-15.

















A zB/ c D E


Figure 3-15. Deformation of an element in an asphalt pavement near the edge of the tire in the
XZ-plane (longitudinal plane).

For a cycle of two passes in two-way loading, an element on the edge of the wheel path in the

asphalt layer will experience a pattern of shear stresses of: A-B-C-D-E E-D-C-B-A. For a

cycle of two passes in one-way loading, an element in the asphalt layer will experience a pattern

of shear stresses of: A-B-C-D-E A-B-C-D-E. The magnitudes of the shear stress are equal and

opposite for positions B and D, as would be expected.

It is not just shear stress reversal, from positive to negative, but the shear stress cycle

reversal from positive to negative to negative to positive (two-way directional loading) versus

positive to negative to positive to negative (one-way directional loading) that may be the cause

for the difference in rut rate and depths between one-way and two-way loading.

Quantification of Shear Stresses Within Asphalt Pavement

To answer the above questions, the stresses that occur within the pavement in the area

thought to be the most critical to the propagation of instability rutting top 2 inches near the tire

edge [Dawley 1990 and Sousa et al. 1991] would need to be quantified. This was accomplished

through three-dimensional finite element modeling of a radial tire contact stresses on the

pavement system used during the FDOT study. The details of the methodology of the finite









element analysis are presented in Chapter 4. The results are briefly summarized here specific to

the HVS testing but were also consistent with the result sin Chapter 4.

The HVS used a super-single radial tire with an average contact stress of 792 kPa (115

psi) and a footprint of 30.5 cm (12 inches) wide by 20.3 cm long (8 inches) (Byron et al. 2003).

Since the exact tire contact stresses were not known for the radial tire used in the HVS, typical

tire contact stress data on a nine-inch wide super-single radial tire rated at 792 kPa (115 psi) was

available and was used in the analysis. Tire contact stresses were based on a tire contact

measurement system developed by M.G. Pottinger [Pottinger 1991], which was especially

developed for tire research, and consists of 1200 distinct measurement points, which register

contact stresses in the x, y and z directions. This resulted in over 3,600 distinct stress

measurements for the nine-inch wide super-single tire radial tire with five ribs and a gross

contact area of 300 cm2 (47 in2) with an inflation pressure of 792 kPa (115 psi). The

measurements provided a high definition of actual tire contact stresses. These stresses were then

applied as the surface loading to the pavement system in the three-dimensional finite element

analysis.

The final three-dimensional mesh consisted of 204,185 nodes-with three degrees of

freedom per node, resulting in a total of 612,555 degrees of freedom. The elements under the

radial contact area had uniform dimensions of 7.62 mm by 10.16 mm (0.3 by 0.4 inches). Table

1 lists the elastic moduli, Poisson's ratio and layer thickness of the pavement layers used in the

analysis.











Table 3-5. Properties of pavement layers used in the finite element analysis.
LAYER Modulus (MPa) Poisson's Ratio Thickness (cm)

Asphalt Concrete 2,071 0.45 10

Base 193 0.45 26

LR Stabilized
128 0.45 30.5
Sub-Base

Subgrade 55 0.45 116



The modulus for the asphalt concrete was determined from IDT resilient modulus tests at

250C on the same unmodified asphalt concrete used in the testing- close to the average ambient

testing temperature in the directional loading analysis [Tia et al 1999]. The base and subgrade

values were based on LBR values conducted by the FDOT and converted to modulus values

using the 1986 AASHTO formula [AASHTO 1986]. The Poisson's ratio was selected to ensure

minimal volumetric changes, to simulate a moving tire load during the shear-driven phase of

instability rutting.

As previously mentioned, the area in the asphalt layer believed most critical to instability

rutting is the upper two inches near the edge of the tire. It was at the location depicted in Figure

3-14, that the shear stresses were culled form the Finite Element Analysis.











Critical/
Location


AC Layer


o8 Critical
Location


Base


00 Y


Subgrade


Figure 3-16. Area of pavement system where shear stresses were obtained.
Figures 3-17, 3-18, and 3-19 display the three directional shear stress magnitudes (YZ,
XY, and XZ planes) in the longitudinal wheel travel directional from the finite element analysis
along the edge of the tire at depths of 1.0, 2.5, 4.0 cms (0.5, 1, and 1.5 inches). Idealized
depictions of the shear deformation for an element in the asphalt layer under the shear
magnitudes reported are presented as well.


Tire Loadr












Direction of Travel


300.0
250.0

200.0
150.0

100.0

50.0

0.0

-50.0

-100.0

-150.0

-200.0

-250.0

-300.0
-3


Figure 3-17. Shear stress in YZ-plane.


-20.0 -10.0 0.0 10.0 20.0

Longitudinal Position, Distance Off Centerline (cm)


Figure 3-18. Shear stress in XZ-plane.


r


-20.0 -10.0 0.0 10.0 20.0

Longitudinal Position, Distance Off Centerline (cm)


BA

A IC


41.0
---2.5
-- 4.0


0.0


300.0

250.0

200.0

150.0

100.0


50.0

0.0

-50.0

-100.0

-150.0

-200.0

-250.0

-300.0
-30.0


-* 1.0
- 2.5
t4.0












300.0
250.0
200.0
150.0
100.0
C 0O.O D1.0
-50.0 C -c2.5
w -100.0---------------------------- -4.0
-150.0
-100.0 D
-150.0
-200.0
-250.0
-300.0
-30.0 -20.0 -10.0 0.0 10.0 20.0 30.0

Longitudinal Position, Distance Off Centerline (cm)


Figure 3-19. Shear stress in XY-plane.

The maximum shear stress in the longitudinal XZ-plane, seen in Figure 3-18, produces

shear stresses about one-half of the maximum shear stress in the transverse YZ-plane, shown in

Figure 3-17. The maximum shear stress in the XY-plane, shown in Figure 3-19, is about one-

third of the maximum shear stress in the transverse YZ-plane, shown in Figure 3-17. It also

interesting to note that the maximum shear stresses occur at depths of 4 cm and shear stresses do

not vary much in the top 4 cm.

To summarize, transverse shear stresses, shear stresses in the YZ-plane, near the edge of

the wheel-path will undergo shearing only in one direction as the wheel approaches and passes --

increasing in magnitude as the wheel approaches and decreasing in magnitude as the wheel

passes as shown above in Table 3-4. This stress pattern of A-B-C C-B-A in two-way loading

and A-B-C A-B-C in one-way loading does not explain the difference in rutting depths

between one-way and two-way loading since the magnitude of B is the same.









However, the stress pattern in the XZ-Plane and XY-Plane from one-way versus two-way

directional loading will vary. One-way loading produces a stress pattern of A-B-C-D-E A-B-C-

D-E while two-way loading produces a stress pattern of A-B-C-D-E-E-D-C-B-A. Since the

stress s pattern in these planes is what differs between one-way and two-way loading, the answer

to why one-way and two-way loading produce different rut rates and depths may lie here.

Analytical Approach

To determine whether the lack of the shear stress cycle reversal, as seen in one-way

loading, has an effect on permanent deformation, a simple analytical model was designed.

Asphalt concrete is a complex material that seldom exhibits solely elastic deformation.

Deformations in asphalt concrete typically contain viscous, visco-elastic and plastic

deformations in addition to the elastic deformations. A simple model that can employ visco-

elastic deformations, including permanent creep deformations, is the Burgers Model, also known

as the four-element model [Ullidtz 1987]. The Burgers Model has the advantages of capturing

instantaneous and delayed elastic deformation responses upon loading as well as instantaneous,

delayed, and permanent deformation responses upon unloading. The Burgers Model is simply a

Maxwell and a Kelvin Model connected in series, as shown in Figure 3-20 [Finedley et al. 1986].





R2r




Il T I
11FB




Figure 3-20. Burgers model.









The constitutive equation for the Burgers Model is described below. Four properties are

required, two elastic constants representing the elastic modulus (R1 and R2) and two viscous

constants (hl and h2). The total strain at time t will be the sum of the strains from the spring and

dashpot in the Maxwell Model and the strain in the Kelvin Model, as written below:

1 = E, 3+ 2 3+ 3

where l8 is the strain in the spring:

G
81 = -
R1

and 82 is the strain in the dashpot, s is the applied stress:

G
82
111

and s3 is the strain in the Kelvin part:

Rl G
3+ 3 =-
12 12

These all can be combined to give the following constitutive equation:


f1i 1i f72 fIfl2 .. fIf2
C + + + 51+ l -l F+ lg
R R2 R2 ) RR2 R2

The simplest way to obtain a solution is to use Laplace Transforms and their inverse. This

has the advantage of simplicity and consistency. Applying the Laplace Transform to the above

equations reduces them to algebraic expressions and transforms them into a function of the

complex variable s instead of time (t) indicated by a caret (A):


&+( < 1 12 )S&+11 2 S 12 11 12^
+ + III+ R+112 = 1,S8 + 1112S 2
R- RR R +2 R1R2 R 2R
1 2 2 12 2









The inverse Laplace transform will bring the above expression back into terms of t.

The Burgers Model was implemented into a MathCAD worksheet, representing asphalt

concrete, to determine whether the reversal of the shear stress pattern (two-way loading) over a

series of two cycles results in differences in total strain as compared to no shear stress pattern

reversal (one-way loading). The finite element stress path evaluation showed that an element in

the upper inch of the asphalt layer does not experience shear stresses until the tire is 20.3 cm (8

inches) away. The maximum stress achieved in the asphalt layer along the edges of the tire in the

longitudinal plane is 103 kPa. Assuming that the 20.3 cm (8 inch) HVS radial tire used in the

testing is traveling at the rated constant speed of 9.6 kph (6 mph), neglecting the deceleration and

acceleration times for slow down and speed up when reversing direction, and the fact that the test

track is thirty feet long, an element would experience a loading for about 0.22 seconds. The

relaxation time between passes for the one-way loading would be 6.6 seconds and for the two-

way loading would be 3.2 seconds.

A sinusoidal load, similar to the actual semi-triangular loading pattern found in the finite

element analysis, with a peak stress of+103 and -103 kPa (+15 psi and -15 psi) was applied for

0.22 seconds. The next loading, representing the next pass, will either be +103 psi then -103 kPa

for the one-way directional loading or -103 and +103 kPa for the two-way directional loading,

each with their respective relaxation times, as shown below.





































I Cycle

Figure 3-21. Longitudinal shear stress pattern in one-way loading.


," 0.22 Seconds 0.22 Seconds

100

75

50

S25
3.2 Seconds
) 0

) -25

-50

-75

-100

-125
1 Cycle 2

Figure 3-22. Longitudinal shear stress pattern in two-way loading.


0.22 Seconds 0.22 Seconds









6.4 Seconds









Properties for the Burgers Model were obtained from IDT creep tests (Roque and Butlar

1992) on a typical fine-grade mix found in Florida similar to the one used in the HVS directional

testing and are presented in Table 3-6.

Table 3-6. Values used in Burgers model to represent asphalt concrete pavement.
Property Value
R1 2.5 Gpa

r1i 86.33 GPaesec
R2 0.5 GPa

712 50 GPaesec



The results from the analysis of the loadings with the Burger Model are displayed in

Table 3-7.

Table 3-7. Results of Burger model under two-way and one-way loading.
Loading Method Strain (mm/mm)
Two-way (Shear Cycle Reversal) 5.434 X 10-s
One-way (No Shear Cycle Reversal) 3.219 X 10-6



The results indicate that the one-way directional loading produces more strain than the

two-way directional loading in the one-dimensional Burger Model. The difference in strain for

the one-way directional loading was two orders of magnitude greater than the two-way

directional loading. Although the magnitude of the predicted deformations is different to that

seen in the field, the differences match the observations that one-way directional loading

produces more deformation than two-way directional loading.

Laboratory tests on stress cycle reversal produce modulus values of one-half to two-thirds

when there is no stress cycle reversal [Kallas 1970]. This suggests that shear stress reversal, even

with greater relaxation times, produces greater deformation as seen in the HVS testing.









Summary and Conclusions

The FDOT testing program conducted focused on the effects of binder modification and

one-way versus two-way loading. However the response of mixtures under-going instability

rutting under these tests was insightful. Analysis of the transverse surface profile and the density-

height changes of the cores before and after testing indicate that the modified sections rutted

primarily due to densification within the wheel-path. The analysis suggests that rutting for the

unmodified sections was due primarily to material being displaced or shoved from the wheel-

path. The lateral shoving and possible expansion of the material once displaced increased the rut

depth measurement and would affect ride performance. Instability rutting therefore is a

combination both densification and lateral shoving, with lateral shoving resulting in higher rut

rates and depths. The modified sections can be said performed better than the unmodified

sections for their ability to resist lateral shoving. A mixtures ability to resist lateral shoving

caused by near-surface shear stresses will determine its resistance instability rutting.

The one-way and two-way loading FDOT study focused on determining which method

would produce rutting in the shortest period of time. The tests results showed that one-way

loading produces greater rut rate and rut depths compared to two-way loading under similar

conditions. The computational Finite Element Analysis, a simple analytical test, and an

understating of the mechanics of the shear stresses occurring within the pavement indicated that

simply viewing rutting as caused by transverse shear stresses is not sufficient. Instability rutting

is truly a three-dimensional phenomenon, a complex problem requiring an investigation that does

not begin overly simplistic. A test that can capture complex stress states and patterns, such as the

Hollow Cylinder test, would be invaluable and an excellent starting point to begin the

understanding of the mechanisms behind instability rutting.









CHAPTER 4
THE APPROXIMATION OF NEAR-SURFACE STRESS STATES IN ASPHALT
CONCRETE THROUGH THREE-DIMENSIONAL FINITE ELEMENT ANALYSES

Background

Finite element analyses have not been extensively used to model three-dimensional tire

loadings due to the complexity of modeling a radial tire in three dimensions and the rather small

(40 60 in2) and highly non-uniform contact area. Moreover, typical pavement structures consist

of a thin layer of asphalt concrete overlying a base course, which rests on the semi-infinite

subgrade. To accurately model the non-uniform loading and provide adequate boundary

conditions requires a large number of elements and the associated large amounts of memory are

not available either on past or current PCs.

Because of these limitations, one approach to model three-dimensional tire contact stress

was to approximate the complex loading conditions with uniform circular loads, use a layered

elastic solution program that solves for uniform circular loads, and apply the superposition

principle [Drakos et al. 2001]. This was done using the computer program BISAR [de Jong et al.

1973]. BISAR allows for circular uniform loads with one stress in the normal direction to the

pavement and another in shear on the pavement at a specific angle. The concept was that a series

of small uniform circular loads of varying vertical and lateral stresses could represent the actual

tire stresses. This produced valuable results demonstrating that radial tires produce high near-

surface shear stresses not found in bias-ply tires. However, these results were based on averaging

numerous measured contact stresses to produce a single uniform circular load. Many of these

uniform vertical loads would then define the radial tire. Furthermore, abrupt changes in

tangential surface stresses on a continuum can lead to infinitely high horizontal stresses affecting

the realism of near-surface stress states [Soon et al., 2003]. A finite element analysis would









provide for more accurate load applications with loads being directly applied as nodal forces and

the use of non-linear elements would provide insurance against superfluous stress concentrations.

The University of Florida recently purchased a Silicon Graphics Interface (SGI) multi-

processor computer, also known as a "super-computer". Because of the extensive memory and

faster computing time than the average PC, the ability to use finite elements in the modeling of

three-dimensional tire-pavement system could be revisited.

In order to better understand the actual stresses that occur in an asphalt pavement from

radial tires the commercial finite element code ADINA [Bathe 2001] was used in modeling the

three-dimensional effects of measured tire contact stresses in a typical pavement configuration.

All pavement layers were assumed to be linear elastic, and dynamic effects were ignored in favor

of promoting a basic understanding of static stress states before complicating the analysis with

dynamic effects. Due to the complicated nature of the measured radial tire contact stresses,

contact surfaces were used extensively to control the size of the problem. The insight gained into

the near-surface stress states believed to be influencing instability rutting from this three-

dimensional analysis would be invaluable.

Three-dimensional Finite Element Model

The three-dimensional finite element model was constructed and analyzed using the

commercial finite element code ADINA. The finite element model consisted of 30,204 27-node

elements of varying dimensions. First, an assessment of the grid size requirements required to

model an attire contact stresses in asphalt pavements was performed. A uniform stress

distribution showed that stresses in the area of interest, near the tire footprint, were not

significantly affected by the extent of the boundaries. The initial assessment demonstrated that

the three-dimensional model should be at least 72 inches deep and extend laterally at least 60









inches in each direction from the center of the tire contact load to adequately represent the semi-

infinite half space conditions associated with pavement problems.

Because of the size of this model, the refinement needed near the tire contact area, and the

desire to remain within the 3:1 element length to width ratio, the resulting memory requirements

for the Silicon Graphics multiprocessor computer available for this analysis exceeded the 1,300

MB RAM memory available. To overcome the limitations associated with building a traditional

mesh, contact surfaces were introduced, where a fine graded mesh representing the loaded

surface was attached ("glued") onto a coarse-graded mesh. This allowed for the introduction of

coarse meshes at distances further away from the loaded area where the change in stress was

more gradual, and far field stresses dominated the response. The use of contact surfaces was

further justified based on the primary area of interest being the near surface area under and

immediately surrounding the loaded tire, thus negating any possible negative numerical effects

of "far away" contact surfaces.

The final three-dimensional mesh consisted of 260,455 nodes with three degrees of

freedom per node, resulting in a total of 781,365 degrees of freedom. Elements were all 27-node

brick elements. Twenty-seven node brick elements were selected over 8-node brick elements in

that 27-node brick elements possess a linear distribution of strain rather than constant strain, are

less stiff, and have a better rate of convergence [Cook 1995]. The elements under the radial

contact area had uniform dimensions of 0.30 by 0.40 inch. Contact surfaces were used for the

transition from the asphalt layer to the base, from the base to the foundation, and from the fine

mesh near the tire contact area to the peripheral areas. The lateral boundaries were fixed in the X

and Y directions, but free in the vertical, Z, direction. The bottom of the mesh was fixed in all

directions. The mesh itself extended 60 inches laterally in each direction and 72 inches vertically.












Figure 4-1 illustrates the final three-dimensional mesh, with Figure 4-2 showing a plan view of


the contact area of the three-dimensional mesh.



z





72 -
64
56
I
N 48
C 40
H 32
E
S 24
16


0 -









Figure 4-1. Three-dimensional finite element mesh used in the analysis.


60 r


30

c


10 F


0 10 20 30
Inches


40 50 60


Figure 4-2. Plan view of the three-dimensional finite element mesh used in the analysis.









Pavement Structure and Loading Conditions Analyzed


Pavement Structure

A typical three-layer (asphalt concrete, base, and subgrade) pavement structure was used in

this analysis [Myers et al. 1999, Drakos et al., 2001]. The asphalt layer had a thickness of 8

inches. The base layer was 12 inches and the subgrade was 52 inches. The properties of each

layer were assumed to be isotropic, homogenous, and linear elastic. Table 4-1 lists the elastic

moduli, Poisson's ratio and layer thickness of the pavement layers used in the analysis.

Table 4-1. Material properties of the various layers used in the analysis.
LAYER Modulus (psi) Poisson's Ratio Thickness (ins)
Asphalt Concrete 100,000 0.45 8
Base 40,000 0.45 12
Subgrade 15,000 0.45 52


A low modulus for the asphalt concrete was selected to correspond to a warm summer

day for a fairly new pavement the most critical time for the onset of instability rutting -- while

base and subgrade values were typical of measured values in the State of Florida. The Poisson's

ratio was selected to ensure minimal volumetric changes, as would be expected from a single

moving tire load.

Loading Conditions Analyzed

The following loading conditions were analyzed:

1. Radial Tires with an inflation pressure of 115 psi
a. 5-ribs with a gross contact area of 48 in2 (R299 tire designation)
b. 4-ribs with a gross contact area of 52 in2 (M844 tire designation)
2. Uniform Vertical Load of 115 psi with a gross contact of 48 in2

Gross contact area for the radial tires refers to the entire area between the outer edges, not

just the area in contact with the pavement. Radial tire stress measurements were obtained from

Smithers Scientific Services, Inc., in Ravenna, Ohio using a method described in Chapter 2.

Stresses and displacements were recorded every 0.20 inch longitudinally (parallel to wheel path)










and every 0.15 inch transversely (perpendicular to wheel path. This resulted in 1,200 stress

locations or 3,600 distinct stress measurements for the radial tire with five ribs and a gross

contact area of 48 in2 and 1,300 stress locations or 3,900 distinct stress measurements for the

radial tire with 4 ribs and a gross contact area 50 in2. The measurements provided a high

definition of actual tire contact stresses. Figures 4-3 and 4-4 depict the variation in vertical

stresses for the two radial tires, and Figures 4-5 and 4-6 display the variation in transverse shear

stress.


1400
1200

1000
800
600


400 12

200


O 3 6 9 12 15 1 21 24
Transverse Dimension (cm)

Figure 4-3. Variation in vertical stress across the 5-rib radial tire used in the analysis.












1400

1200

1000


800o


0 3 6 9 12 15 18 21 24 27 30


13vI
Jo Q


Transverse Dimension (cm)

Figure 4-4. Variation in vertical stress across 4-rib radial tire used in the analysis.




50











15


0 3 6 9 12 15 18 21 24 0
C
h-
Transverse Dimension (cm)


Figure 4-5. Variation in horizontal stress across 5-rib radial tire used in the analysis.





















0


-25 3
0 9 12 15 i 21 27 30

Transverse Dimension (cm)

Figure 4-6. Variation in horizontal stress across 4-rib radial tire used in the analysis.

The figures display the truly non-uniform loading conditions and the variation in stress

intensity not only across the tires but also across each individual rib.

Load Application

The measured tire contact measurements reported above were uniform stresses acting over

areas of 0.03 in2 in size. Three different stresses were provided, namely vertical normal stresses,

transverse shear stresses, and longitudinal shear stresses. Ideally, each uniform stress should be

applied to a single element. Unfortunately, the number of elements needed would have again

exceeded the memory of the Silicon Graphics multiprocessor computer available for the analysis.

Thus, the use of fewer elements was required under the contact area, which subsequently

required the determination of the equivalent nodal forces to be applied to each node. The

appropriate nodal forces for each element were determined by converting each uniform stress

into an equivalent concentrated force. These forces were then applied to each element and

distributed over the element nodes according to the rules described below [Hughes 1987].









If it is assumed that the concentrated forces {/} has componentsf,,fy, andf, then the

element load vector, {re}, acting on a surface of an element, is defined by:
{e} L fJ[N]T}IdSls
Se

where [N]T is the transpose of the shape function matrix, and {f} is the surface traction

vector. The contribution of {f} to {re} can be determined by viewing the concentrated force as a

large traction, {(}, acting over a small area, dS. Subsequently, the concentrated force vector {f

} can be denoted as:

{f} = {}*dS

The integral of [N] ({}dS thus becomes [N] {f}, resulting in Equation (1), with n

concentrated forces, becoming:

n T
{re} = [N] {f


where [N], is the value of [N] at the location of {f}i.

Elements under the loaded area were selected to be 0.30 inch by 0.40 inch to meet the

memory requirements. Each element thus had 24 (eight x, y, and z forces) point forces acting on

it. These point loads were then converted to the appropriate nodal forces through the method

described above. Finally, Figure 4-7 shows a plan view of the radial tire contact stresses applied

as nodal forces onto the pavement model surface at locations consistent with five tire ribs.

Figure 4-8 shows a cross-sectional profile of the tire contact stresses.

























Figure 4-7. Contact area and the radial tire nodal forces used in the pavement response analysis.


0.0 Inches 9.3

Figure 4-8. Cross-sectional view of applied tire nodal forces used in the pavement response
analysis.

Axisymmetric model

A two-dimensional axisymmetric model (Figure 4-9) was generated in ADINA to provide

a comparison between the stresses induced by a circular uniform vertical load and the more

complicated radial tire loading effects. Because of the symmetric nature of the problem, only

one half of the loaded area is modeled. The finite element model is 72 inches tall and 30 inches

wide, with a uniform vertical surface load of 115 psi, distributed over a radius of 4 inches. The

elements used consist of 8-noded isoparametric elements, with 72 vertical rows of elements, each

containing 99 elements, for a total of 7128 elements. The layer thicknesses and elastic properties

are the same as in the three-dimensional finite element model, discussed previously.











72 r-


N
40
C
H
E 32


24


16





0 -

Figure 4-9. A cross-sectional view of the axisymmetric finite element mesh used for comparison
purposes.

To evaluate the results from the axisymmetric finite element model, a comparison was

performed between predicted shear stresses at the edge of the loaded area and those obtained

from a semi-analytic layered-elastic theory solution, using the program BISAR [de Jong et al.

1973]. Figure 4-10 shows that the shear stress predictions obtained with ADINA [Bathe 2001]

and BISAR [de Jong et al. 1973] are very similar for the exact same loading conditions, meaning

that the axisymmetric model adequately captures the loading response due to the circular

uniformly loaded vertical load.










Shear Stress (psi)
0.0 5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0 45.0
0.0

0.5
-A- BISAR
10 -0-ADINA

1.5

2.0

2.5

3.0

3.5

Figure 4-10. Comparison of shear stresses under the edge of a two-dimensional axisymmetric
circular uniform vertical load predicted with BISAR and ADINA.

Three-Dimensional Solution Process

There are four different types of solution schemes available in ADINA, namely, 1) direct

solver, 2) sparse solver for very large problems, 3) iterative solver, for nonlinear problems, and

4) multigrid solver, for parallel processing solutions. The direct solver requires a large amount

of storage and is not recommended for large three-dimensional models [Bathe 2001]. The

multigrid solver is intended for large three-dimensional problems with very large systems of

equations, but when contact surfaces are used, ADINA does not allow for the use of the

multigrid solver. The best solver for large memory limited problems is the sparse solver.

Unfortunately, for the number of equations anticipated, the system requires a 64-bit version

solution solver, with only the 32-bit version currently available for this research. Hence, since

the iterative solver is also recommended for large problems, it was used by default.

The equilibrium equations to be solved in a non-linear static analysis of a finite element

model with contact surfaces in ADINA are:

t+AtR- t+tF = 0









where t+AR is the vector of the external nodal loads and t+tF is the force vector equivalent

to the element stresses at time t + At [Bathe, 2001]. In non-linear analysis three iteration

methods/schemes are available in ADINA, namely, 1) full Newton method, with or without line

searches, 2) modified Newton method, with and without line searches, and 3) the Broyden-

Flecther-Goldfarb-Shanno (BFGS) matrix update method. In this study, the full Newton method

with lines searches was employed, based on its ability to converge and obtain accurate solutions

[Bathe 2001]. The iterative solution process required 1295 Mbytes of RAM memory, with a

resulting solution time of 9,000 seconds, using a single processor.

Results of Three-Dimensional Analysis

The vales to be reported from the FEM analysis should be viewed as more "qualitative"

than quantitative. The reason for this approach is that the asphalt pavement was modeled as a

linear elastic material and the FEM analysis was not as robust in terms of the number and sizing

of elements as was originally hoped. However, the results should provide insight into the stress

states that occur within the asphalt pavement from radial tires that will help lead to a better

understanding of the mechanisms behind instability rutting.

As an initial evaluation of the three-dimensional stress states that occur within the

pavement in areas where instability rutting is observed was undertaken. Stress states in the

longitudinal direction direction of tire travel -- were analyzed to compare the three-dimensional

differences between radial tire loadings and uniform vertical loading. The location selected for a

stress-path comparative analysis was a point in the asphalt pavement one (1) inch away from the

loaded area and one (1) inch deep into the pavement. The stress states at this point was also

studied 10 inches ahead and behind in the longitudinal direction (x-direction) of the loaded areas.









Outer
Rib 25 mm

25 mm

25 mm







x
z

Y
Y
Figure 4-11. Location within asphalt pavement from finite element analysis for preliminary
stress state evaluation.

The maximum shear stress, the confinement stress, and the orientation of the principal

stresses were analyzed. The maximum shear stress, Tmax, is defined as the difference between the

first and third (largest and smallest principal stresses in magnitude, respectively) divided by two:

zmax = (51 03)/2

Confinement, p, is defined as the summation of the three principal stresses divided by three:

p = (C + C2 + G3)/3

Presentation of the orientation of the principal planes was done by calculating the Lode

Angle. The Lode Angle, 0, indicates the orientation of the line from a three dimensional stress

state from the hydrostatic axis. The Load Angle varies from +300 to -300. The +300 state

indicates the stress sates are the principal stresses, with Io > C2 = 03, which can be viewed as a

state of triaxial compression. The -30 state indicates the stress sates are the principal stresses,

with c1 = C2 > 03, which can be viewed as a state of triaxial extension. A Lode Angle of 00 would

indicate a state of hydrostatic stress, i.e. CI = C2 = 03. The Lode Angle was selected as an

efficient means to depict the rotation of the principal planes in a complex three-dimensional










problem and to highlight the differences between radial tire loadings and uniform loadings

conditions.

The following four (4) figures depict the maximum shear, the confinement, the stress-

path, and Lode Angle in the longitudinal direction. An "image" of the outer loaded area (length

of the loaded area is 6 inches) is depicted on the y-axis of the graphs to help visualize the

position of the stress magnitude compared to the tire position. Although these stress states are

from a static analysis, a stress path of the loading from a moving tire load also can be inferred

from the images.


0 5 10 15 20 25 30 35
MAXIMUM SHEAR STRESS (PSI)

Figure 4-12. Maximum shear stress in longitudinal direction along outer tire rib.
























4






I -2

0
0 ----0_0-P3^
o 2










-10
0 5 10 15 20 25
CONFINEMENT (PSI)


Figure 4-13. Confinement stress in longitudinal direction along outer tire rib.


40



35



30 O Uniform Load



u 25



En 20 2 0"5
E4 o

'm 15



10
5 -














10








0 5 10 15 20 25 30

Confinement (psi)

Figure 4-14. Confinement and maximum shear stress path.
i25-----^-----








I 15------.----^0-------














Figure 4-14. Confinement and maximum shear stress path.


















?u 6

S 4

-3 0 -2 200 -5 1 -5 00100 5
z
2

0

2- -

I -4j

-6

-8

-10
-300 -250 -200 -150 -100 -50 00 50 100 150 200 250 300
LODE ANGLE (Degrees)

Figure 4-15. Lode angle in longitudinal direction along outer tire rib.

The non-uniform stresses caused by the radial tire's configuration (a 5-rib M844 radial


tire) induce approximately 40% more shear stress within the pavement than the uniform loaded


area. The largest maximum shear stress occurs directly adjacent to the middle of the loaded area.


In addition, the magnitude of the confinement of the radial tire is approximately 10% less than


that of the uniform loaded area. This is due to the fact that the radial tire's vertical load is


distributed unevenly more vertical load is concentrated in the middle ribs and less at the outer


edges.


The Lode Angle orientation suggests that the lateral stresses induced by radial tire


configuration produce more of an extension state of stress than a vertical uniform load. Since a


Lode Angle of -300 was not obtained, a simple triaxial extension test will not be able to capture


the principal plane orientation that is induced by radial tires.












As was somewhat expected, the radial tire produced much higher ratios of shear to


confinement than the uniform load condition. This analysis was conducted on linear elastic


material; if a material with failure/yield criteria with pressure dependency was instead used to


model the asphalt pavement, radial tires may induce stress states into the pavement that will


result in stresses that may yield the material and will result in a re-distribution of stresses that are


not predicted by a simplified uniform vertical load. It is important to note that the largest ratio for


the radial tire occurs at the corner of the loaded area rather than adjacent to the middle of the tire


where the maximum shear was calculated.


Finally, the three directional shear stresses (Txy, Tyz, Txz) were compared to evaluate the


contribution of each to the maximum shear stress. The shear stresses at the locations analyzed


above are presented below in the following figure.


6 SHEAR YZ
SSHEAR XZ
II 4




o 0
01


8
O
I






















rib.
-- 4
--
a 2





- 8














rib.
-J 6

-8







rib.









The largest shear stress occurs within the transverse plane the YZ-plane, which is where

the maximum shear also occurs. This suggests the most critical plane for investigating instability

rutting is the transverse plane adjacent to the center of the loaded area. Transverse shear stresses

have generally been thought to be the primary mechanisms behind instability rutting [Drakos et

al. 2001, Dawley et al. 1990]. However, reversal pattern of the other directional shear stresses,

the ratio of maximum shear to confinement, and the results of the one-way and two-way loading

studies, indicate that instability rutting is a true three-dimensional phenomenon where shear

stresses in the non-transverse plane play a role. To simplify the further investigation, the research

focused on the transverse plane as the critical plane which would turn the complex phenomenon

into a two-dimensional study. The three-dimensional aspects of instability rutting can not be

ignored.

Maximum shear stresses within the transverse plane are presented along the middle

transverse cut of the 5-rib radial tire, 4-rib radial tire, and uniform vertical loading conditions,

respectively with directional arrows showing the direction of the smallest angle formed between

the maximum shear stress and the horizontal.



S350
300
250
-4 FT k -%-- -
1200
150

0 5 10 15 20 25 30 35 50

Later Location (cm)

Figure 4-17. Maximum shear stress magnitude (in kPa) and direction under the 5-rib radial tire
loading condition.











0 350

-2 2 1 1 4 4 250
-4 1"200
l c t
-6 150
_100
50
0 5 10 15 20 25 30 35 40 45 -
Lateral Location (cm)

Figure 4-18. Maximum shear stress magnitude (in kPa) and direction under the 4-rib radial tire
loading condition.




0 =350
/2 j l ai300

200

L0L 100
0 5 10 15 20 25 30 35 50
Lateral Location (cm)

Figure 4-19. Maximum shear stress magnitude (in kPa) and direction under the uniform loading
condition.

A further investigation of stress states in this area revealed that there existed two critical

locations were the maximum shear stress was the highest within the pavement area adjacent to

the tire loading.



Outer Tire Rib


A 0 .3"
1.15"
0.5"
Be

1.375"


Figure 4-20. Critical stress locations for high shear stress.










The following figures present through Mohr Circles the stress state at these critical

locations for the three loading conditions analyzed.


60
Horizontal Plane
Stresses
Vertical Plane
40 Stresses
M844

20 R299




0
\\\ Uniform




-20 0 20 0 \ 80 100 120 140
2- = 56 psi

-20

\ =-15psi

-40

3 = 52 psi
03=-12 psi
-60-


Figure 4-21. Stress states at point A for the three loading conditions analyzed.















40
Uniform

R299
20
M844


0-
-20 0 20 40 0 80 100 120 16 0


-20

c1 = 65 psi

-40

01 = 60 psi
03 =-8psi \
\ 03 = -12psi
-60


Figure 4-22. Stress state at point B for the three loading conditions analyzed.

The figure clearly indicates that the normal and shear stresses induced by radial tires are

driving the stress state into planes of tension compared to a uniform load. The figures also

display that the principle stresses are in a state of extension -- with extension defined as the first

principle stress having an angle of less than 450 with the horizontal plane. A compression state of

stress is defined as the first principle stress having an angle equal to or greater than 450 with the

horizontal plane [Lambe and Whitman 1969].

Significance of High Shear at Low Confinement

Asphalt mixtures are known to be pressure dependent materials and have been modeled as

Mohr-Coulomb materials [Krutz and Sebaaly 1993, McLeod 1950]. This is significant because

HMA mixture strength parameters are often obtained at high these higher confinements. The c










and ) parameters when idealizing asphalt mixtures to a Mohr-Coulomb envelope are typically

based on strength tests in these high confinement regions and extrapolated, but not measured, to

areas of low confinement. At elevated temperatures, asphalt mixtures begin to behave like

granular materials with little cohesion as the binder becomes less viscous. In addition, the

envelope in these low confinement regions may be curved, as indicated by the dashed line in

Figure 4-23, which is sometimes the case for granular materials [Lambe and Whitman 1969].


25
Asphalt Concrete at -
20 Normal Operating ..-
15 Temperatures -
15 Temeraure Granular Material
10
SAsphalt Concrete at
5 High Temperatures
0
-5

-10

-15

-20

-25
-20 -10 0 10 20 30 40
Normal Stress

Figure 4-23. Typical strength envelopes in hot mix asphalt and granular materials.

Even unconfined vertical loading conditions will result in a Mohr-Coulomb failure circle

with one principle stress at 0, meaning material response is determined from stress states and

principal plane orientations that do not appear representative of the stress conditions in the

asphalt occurring from radial tire loadings. In terms of rutting, the material response in areas

thought to be the critical to instability rutting are under high shear and low confinement.









Summary and Conclusions

The shear stresses of the radial tires differ from the uniform load in both magnitude (up to

20% greater) and distribution. The stress distribution also differs between the radial tires. All

three loading conditions indicate the formation of shear planes under the loaded area. The shear

planes will tend to shove the material up and away from the tire resulting in rutting within the

asphalt pavement. The shear stresses are more prevalent under the radial tires near the surface

compared to the uniform loading.

The three-dimensional analysis of measured radial tire contact stresses indicated that stress

states in the asphalt layer are characterized by low levels of confinement, and even tension,

coupled with high shear stress. In contrast to this, stress states induced by uniform vertical

loading conditions, which are traditionally assumed in pavement analysis, are characterized by

higher levels of confinement and lower shear. Furthermore, these more critical stress states from

radial tire contact stresses occur in the near surface region, near the edge of the tire, where

instability rutting is known to occur. Uniform vertical loadings do not produce these critical

stress states in this region.

The results presented also imply that the characterization of instability rutting requires

testing at these low confinement (and sometimes tensile) stress states and principal lane

orientation, rather than at the higher stress states typically used in the strength characterization of

mixtures.

Accurate modeling of radial tire contact stresses is necessary to describe the mechanism of

instability rutting. It appears that an accurate description of mixture properties (shear strength

and deformation responses) at stress states (high shear at low confinement) in the critical

instability-rutting region (edge of the rib at shallow depths) may be necessary to properly

evaluate a mixture's instability rutting potential. The next chapter investigates what laboratory









tests would be available to produce similar stresses and orientations that were indicated in the

FEM analysis as occurring in the asphalt pavement in these critical regions.









CHAPTER 5
CREATION OF A HOLLOW CYLINDER DEVICE

Introduction

The computational analyses presented in Chapter 4 found radial tiers induce higher shear

stresses at lower confinements than bias-ply tiers in areas of the asphalt pavement where

instability rutting is observed to occur. It is believed that instability rutting is driven by high

shear stresses at low confinements. The orientation of the principal planes induced by tire loads

in these areas is not aligned with the vertical and horizontal planes, but orientated 20-30 from

the horizontal. If these stress conditions are critical to instability rutting then it is of interest to

test under these critical stress conditions. Testing under these conditions is believed will help

determine a mixture's susceptibility to instability rutting more accurately than current tests which

do not test under these conditions.

The laboratory equipment currently available triaxial device cannot obtain these critical

stress states orientation of the principal planes. Currently, two laboratory testing devices can

match the stress states mentioned the true triaxial device and the hollow cylinder device. The

true triaxial device is a complex, rare, research-only tool that has no feasible industrial or

mainstream research capability. The hollow cylinder device, on the other hand, can be instituted

with typical research testing equipment. Although, hollow cylinder testing is not typically used

outside of research institutions it is more widely understood, recognized, and used at research

institutes far more than the true triaxial device. For these reasons, a hollow cylinder testing

device (HCTD) was constructed at the University of Florida to test asphalt mixtures under these

critical stress states. This chapter describes the development of the HCTD at the University of

Florida.









Hollow Cylinder Specimen Dimensions

The first task prior to developing the mechanical apparatus was to determine the

dimensions of the hollow asphalt concrete (AC) specimen. Once the specimen dimensions were

determined, the fabrication of the HCTD could proceed and an analysis of stress concentrations

under typical testing conditions for the selected geometry could be undertaken.

The dimensions of the hollow cylinder will affect the stress distribution across the

specimen under axial and torsional loading. The wall thickness must be thick enough to

accommodate the range of aggregate sizes in the asphalt mixture to ensure a representative

sample but, be thin enough to allow for more uniform stress distribution of across the wall. In

addition, means and methods to easily and readily construct the hollow cylinder must also be

considered.

A previous study investigating rutting induced by fighter jets on asphalt concrete

runways employed a one-inch thick wall thickness [Crockford]. This study employed cyclic

torsional testing with inner and outer pressure differentials. Present researchers also use a one-

inch thick wall thickness in their HCTD for the investigation of low-temperature crack

propagation in asphalt pavements [Buttlar] these studies do not employ torsion on the samples

but a greater inner pressure than outer pressure to induce a tangential "hoop" stress.

In Florida, SP-9.5mm and SP-12.5mm are typically the most common Superave asphalt

structural mixes employed. These size mixtures would therefore be the most likely employed in

pavement design and studied in pavement analyses. The largest aggregate size that would

therefore be encountered with some frequency for these mixture designations is the nominal

maximum aggregate size of 9.5 or 12.5 mm (a SP-9.5mm mixture has a maximum aggregate size

of 12.5 mm and a SP-12.5mm has a maximum aggregate size of 19 mm.) The HCTD wall









dimensions must be sized to ensure that a representative sample of a SP-9.5mm to a SP-12.5mm

mix is present.

The University of Florida employs a ServoPac Gyratory compaction device for creating

laboratory asphalt concrete specimens. The ServoPac can produce either 100mm or 150mm

diameter specimens. These diameter sizes are common for gyratory compactors used in asphalt

research and industry. Specimens made to a 100mm diameter can be up to 150mm in height.

Specimens made to a 150mm in diameter can be up to 135mm in height.

Most guidelines established for hollow cylinder wall thicknesses have been based on soil

testing. For instance, as a rule of thumb, it is recommended that the wall thickness should be 10

to 25 times the average grain size to ensure a representative [Saada and Townsend]. If one used

this recommendation for asphalt concrete testing, the typical SP-12.5mm mixture has an average

grain size (defined as D50) of approximately 4mm, thus a wall thickness ranging from 40 mm to

100 mm would be recommended. (The largest particle size in a SP-12.5mm is 19mm.)

Based on the above information, the outer diameter of the hollow cylinder specimens will

either be 100mm or 150mm and the height either 150mm or 125mm for the respective heights.

The inner diameter is the one parameter that must be designed for that best meets the desires of a

representative sample and stress uniformity. An initial investigation looked at inner diameters in

25mm increments that would produce wall thicknesses of 25mm or greater. Table 5-1 indicates

the possible specimen dimensions for the hollow cylinder.











Table 5-1. Possible hollow cylinder specimen dimensions.
Specimen Outer Diameter, do (mm)
Parameter
100 150
Height, L (mm) 150 135
Inner Diameter, di (mm) 25 or 50 50, 75, or 100
Wall Thickness (mm) 37.5 or 25 50, 37.5, or 25
di/do 0.25 or 0.50 0.33, 0.50 or 0.67
L/do 1.5 0.9



As mentioned in Chapter 2, there are guidelines, recommendations, and techniques that are

available that can be employed to evaluate a hollow cylinder specimen. These are in place to

help ensure that the hollow cylinder has the least amount of stress non-linearity and can be

considered as "representative" sample. An analysis based on closed-form solutions was carried

out to determine the effects that the different proposed sample geometries have on the

distribution of the tangential, radial and shear stress across the wall of the proposed hollow

cylinder geometries. Based on this analysis, the selection of geometry could be made based on

knowledge of stress distributions, limitations, and deficiencies between the different geometries.

A finite element analysis would then be employed to investigate end effects and stress

concentrations on the proposed hollow cylinder geometry not possible from the closed-form

solutions. A determination based on the quality of the results from the FE analysis would

determine whether the geometry could be implemented for design and production.

Closed-form analysis of stresses across hollow cylinder wall

The level of stress non-uniformity of the radial and tangential stresses across a hollow

cylinder specimen due to different internal and external pressures was undertaken. Each of the

seven possible specimen geometries had an inner and outer pressure applied to achieve an









average uniform radial stress, or, of 20 psi and an average uniform tangential stress, ot, of 30 psi.

The average stress and the inner and outer pressures stresses were calculated from equations

found in Chapter 2.

Table 5-2. Inner and outer pressures applied to proposed hollow cylinder geometries.
di (mm) do (mm) di/do Inner Wall Outer Wall
Pressure, pi (psi) Pressure, po (psi)
25 100 0.25 5.0 23.8
50 100 0.50 15.0 22.5
50 150 0.33 10.0 23.3
75 150 0.50 15.0 22.5
100 150 0.67 17.5 21.7



The quantitative values of the inner and outer diameters do not play a role in this analysis,

only the ratio of inner diameter to outer diameter (di/do). Figures 5-2 and 5-3 display the non-

linearity of the radial and tangential stress distributions across the wall of the hollow cylinder

specimen for the different specimen geometries (note that r is defined as the radius from the

center minus the inner radius, ri). The stresses have been normalized, so a value of one represents

the average stress, either 20 psi for the radial stress or 30 psi for the tangential stress.















1.2


1


0.8


L:' 0.6
ID
0.4


0.2


0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1
r/(ro-ri)

Figure 5-1. Ratio of radial stress to average radial stress across the wall of the specimen for
various geometries.


1.6

1.4 -*-di/do= 0.25
1.4
*- di/do = 0.33
1.2 di/do=0.50
h tdi/do = 0.67
1

0.8

13 0.6

0.4

0.2

0
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1
r/(ro-r,)

Figure 5-2. Ratio of tangential stress to average tangential stress across the wall of the specimen
for various geometries.

When a torsional stress is applied to a hollow cylinder specimen, the shear stress varies


linearly across the wall thickness. The maximum shear is on the outside edge, while the


-k tdi/do = 0.25
r tdi/do = 0.33

r tdi/do = 0.50

tdi/do = 0.67










minimum shear is on the inside edge. Each specimen was evaluated by determining the slope of

the shear across the wall when a torsional stress is applied. Figure 5-4 displays these results.

Again, the shear stress has been normalized based on the average shear stress occurring on the

specimen based on Equations from Chapter 2.


1.60

1.40

1.20 ---

r 1.00

0.80 0 _*-di/do = 0.25

P 0.60* -+ di/do = 0.33
0 di/do = 0.50
0.40
-di/do = 0.67
0.20

0.00
0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0
r/(ro-ri)
Figure 5-3. Ratio of shear stress to average shear stress across the wall of the specimen for
various geometries.

Table 5-3 gives the stress non-uniformity coefficient 33 from Chapter 2 for shear stress

across the specimen for the different possible geometries. The values reported parenthetically are

the percentage above the recommended value of 0.11 [Vaid et al. 1983].

Table 5-3. Beta 3 shear stress non-uniformity quantities for the various geometries.
di/do 33
0.25 0.28 (155%)
0.33 0.24 (118%)
0.50 0.16(45%)
0.67 0.09



The results from Figures 5-2 to 5-4 indicate that a thinner wall produces a greater

uniformity of stress across the hollow specimen than a thick wall, as would be expected. This









difference is markedly seen when comparing either a di/do of 0.25 or 0.33 to 0.67, but not as

great when comparing a di/do of 0.50 to 0.67 for the radial and tangential stress distribution.

Thus, a di/do of 0.67 is not seen as a great improvement on a di/do of 0.50 for radial and

tangential stress distribution. The 33 parameter is significantly affected by wall thickness.

At this point a decision of the geometry of the hollow cylinder specimen became clearer,

but still engineering judgment would have to be employed. The specimens with outer diameters

of 100mm were excluded for the following reasons:

1. An ID of 25 mm resulted in too much stress non-uniformity

2. An ID of 50mm (25mm wall thickness) would produce a average grain size ratio of an SP-
12.5 to wall thickness below 10

3. Latex membranes to seal the inner radius of a 100mm hollow cylinder are not readily
available (The outside of the hollow cylinder as well as the inside of the specimen (pi and
po) could potentially be under different pressures. Water or air could be used as the
pressuring medium; if water is used then asphalt specimen must be protected against water
infiltration.)

Specimens with an outer diameter of 150mm would have to have an inner diameter

approximately between 75mm to 100mm. The improvement in stress uniformity for radial and

tangential stress is not significant, but for shear it is. However, a 25mm wall thickness was

considered more of a concern lack of a representative sample and grain size to wall thickness

ratio -- than a high 33.

Based on the above information, desire for a representative sample, and engineering

judgment, a specimen size with an outer diameter (OD) of 150 mm, an inner diameter (ID) of

70mm, and a height of 135mm was selected. The specimen size would provide a wall thickness

of 40mm meeting the recommendations for hollowing cylinder testing of soils by Saada for an

average SP-12.5mm mixture and helping ensure a more representative sample. Figure 5-5

displays the specimen dimensions selected that will be used for the HCTD:















ID = 70mm


OD = 150mm











Height = 135 mm


Figure 5-4. Hollow cylinder specimen dimensions.

A finite element analysis was now employed to investigate end effects and stress

concentrations on the proposed hollow cylinder geometry not possible from the closed-form

solutions. A determination based on the quality of the results from the FE analysis would

determine whether the geometry could be implemented for design and production.











Finite element analysis of stress non-linearity and concentrations

The purpose of the finite element analysis was to detect and evaluate stress concentrations

that occur for the selected geometry as well for any coring eccentricities that may occur. The

finite element program ADINA [Bathe 2001] was employed to model the different specimen

geometries under different loading conditions. The analysis employed a linear elastic material

model with a modulus of 100,000 psi and a Poisson's ratio of 0.35. The mesh consisted of 832

twenty-seven-node brick elements. The bottom boundary was fixed in all directions. The top

boundary was free to move only in the z-direction for the radial/tangential and vertical stress

analysis. For the torsional stress analysis, the top of the specimen was kept in plane and a

displacement was prescribed resulting in 0.01% strain. The analysis was three-dimensional.


















Figure 5-5. Three-dimensional finite element model of the hollow cylinder specimen for end
effects analysis.

Again, an outer pressure of 22.5 psi and an inner pressure of 15 psi was placed on the

specimen. This, according to the closed-form solution, would produce an average stress of 20 psi

in the radial direction and 30 psi in tangential direction. The closed form solution predicts a









maximum radial stress equal to the outer and inner pressures at the associated boundaries, i.e., a

maximum radial stress of 22.5 psi on the outer wall and 15 psi on the inner wall. The closed form

solution predicts a maximum tangential stress of 35 psi on the inside wall and a minimum

tangential stress of 27.5 psi on the outside wall.


-22
1-14
-6
2
S10
18
26
32


Figure 5-6. Tangential (hoop) stress distribution across hollow cylinder wall from finite element
analysis.




















12
-6
0
6
12
-24










Figure 5-7. Radial stress distribution across hollow cylinder wall from finite element analysis.

Both figures display stress concentrations near the ends of the hollow cylinder specimen

from the loading condition. It appears that the middle third is relatively little affected by the end

restraints for the radial stresses, but the tangential stress actually is uniform. The tangential stress

distribution along the middle third, although affected by the end constraints in that it doesn't

match the closed-form solution, may actually be considered an improvement since it's now

appears uniform.

For a closer inspection of the analysis, the stress distribution form the closed-form analysis

was compared across the wall at three locations from the finite element analysis. Figure 5-8

depicts the heights along the hollow cylinder where the stress distribution across the wall was

compared to the closed-form solution.












22.5 m


n


45 mm




67.5 mm


I I V
Figure 5-8. Three vertical locations where stress distributions across the wall from the finite
element solution were compared to the closed-form solution.


40.00


35.00


30.00


1 25.00
b
20.00


15.00


10.00


0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1
r/(ro-ri)

Figure 5-9. Comparison of the tangential stress across the hollow cylinder wall from the closed-
form solution and the finite element analysis with end constraints.


135 mm


ADINA --45mm

-ADINA -- 22.5 mm
-^ADINA 22.5 mm










24.00

22.00

20.00

18.00 -
S-*-Closed Form
b 16.00
t --ADINA -- 67.5mm

14.00 ADINA--45mm

-ADINA -- 22.5 mm
12.00

10.00
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1
r/(ro-ri)

Figure 5-10. Comparison of the radial stress across the hollow cylinder wall from the closed-
form solution and the finite element analysis with end constraints.

From Figure 5-9, the middle 45mm appears to have a relatively uniform tangential stress

distribution. The uniform tangential stress is near to the average tangential stress of 30 psi

calculated form the closed-form solution. From Figure 5-10, the middle third and some more

closely matches the closed-form solution. The end constraints caused stress concentrations only

in top and bottom third as shown by Figures 5-9 and 5-10. Thus, an ID of 70 mm on a 150mm

OD sample can be considered acceptable in terms of stress distribution.

The three-dimensional results in Chapter 4 indicate that the normal longitudinal stress (the

second principal stress) is approximately 60% the transverse normal stress (the first principal

stress). As previously stated in Chapter 4, the transverse plane within the pavement next adjacent

to the loaded area is assumed to be the critical plane for instability rutting. Instability rutting is

primarily the effect of stresses within the transverse plane both shear and normal stresses

acting within the transverse plane reducing the problem effectively to two-dimensions. In

addition, testing equipment limitations did not allow for testing under different inner and outer









pressures at the time of this work, requiring the inner and outer pressures remaining the same

during testing -- thus eliminating a "hoop" stress being induced in the specimen and reducing

the problem to twp-dimensions. Based on the previous analyses, applying no differential pressure

between the inner and outer pressures will only reduce stress concentration and stress

discontinuities across the specimen creating a more uniform stress within the specimen. Future

testing that employs differential inner and outer pressures will likely require additional

investigation into stress discontinuity due to loading conditions.

Creation of a Hollow Asphalt Concrete Specimen

Hollow asphalt concrete specimens can be created through methods that require unique

molds, methods, and machining [Sousa 1988]. The selection of the University of Florida hollow

cylinder was based a desire to ensure that specimens could be easily and quickly produced with

existing production methods and machinery. The method selected, that is both effective and

efficient, is to core out the center of a laboratory compacted solid cylindrical specimen (or a field

core). Asphalt concrete cylinders are easily and routinely created using gyratory compactors. A

coring drill is also available. Coring bits typically come in quarter-inch diameter increments

producing a diameter (inner diameter of hollow specimen) of an additional four-hundredth of an

inch. Currently, another researcher employs this method for the production of hollow cylinders

[Buttlar 1998]. Figure 5-11 depicts the hollow cylinder specimen production concept.


















Solid Cylindrical
Asphalt Concrete


Core Out Center
with Drill


Hollow Cylindrical Asphalt
Concrete Specimen and Solid
Core


Figure 5-11. Hollow cylinder specimen production.

It has been well documented that gyratory compacted samples do not have uniform

density. Specimens compacted in the gyratory compactor have higher density (lower air voids) in

the center and lower density (higher air voids) near the top and bottom and along the cylinder

edges. This is more pronounced in taller specimens and wider specimens [Harvey et al. 1994,

Masad et al. 1999, Tashman et al. 2001].

Other researchers who employ coring for hollow cylinder production have not found

density non-uniformity to be such an issue [Buttlar 1998]. Nonetheless, air voids of the solid

cylindrical specimen prior to coring and the post-cored solid center can be calculated using

ASTM D1 189 and D2726. From these quantities, a reasonable estimate of the air void in the

hollow specimen can be calculated through volumetric averaging.

Hollow Cylinder Specimen Production

Aggregate Preparation and Batching

The first step in production is to obtain and procure the aggregate as described below:

* Virgin aggregate obtained from the field is dried in the oven at a temperature of 300F for
twelve hours and then allowed to cool at room temperature until it can be handled.

* The virgin aggregate is then sieved and separated into individual particle sizes based on
Superpave sieve sizes 34" (19 mm), 12" (12.5mm), 3/8" (9.5mm), #4, #8, #16, #30, #50,
#100, #200, and pan (minus #200).









* The aggregate material is then batched according to the Job Mix Formula (JMF) into
5000g samples.

Mixing

The next step in specimen production is to combine the aggregate with the asphalt binder is

a process known as mixing described below:

* The batched aggregate (5000g), asphalt binder, and mixing equipment (pans, buckets,
mixing tools) are placed in an oven for two hours at 3000F.

* The appropriate amount of asphalt binder based on the Superpave design is added to the
batched aggregate and thoroughly mixed until the aggregate has been fully coated with the
binder.

Short Term Oven Aging (STOA) and Compacting

Once mixing is complete, the material is aged and compacted as described below:

* The mixed asphalt concrete is spread evenly in a shallow pan and placed in an oven at
2750F for two hours. After one hour, the mixture is stirred to ensure uniform aging. This
process is known as STOA.

* After the STOA, the mixture is then placed in the 150mm diameter Servopac gyratory
mold (which as been heated at 2750F for two hours) and compacted at 1.250 gyratory angle
with 600kPa ram pressure for the appropriate number of gyrations to achieve 7% air voids.

* The compacted cylindrical specimens are extruded from the mold and allow to cool for 24
hours.

* The cylindrical specimens are then sawed on each side. The height of the cylindrical
sample after sawing is 135mm.

* The Bulk Specific Gravity (GMB) of the sawed specimens was then taken in accordance to
ASTM D1189 and D2726.

* The percent air voids was then calculated from the GMB and the GMM of the specific
mixture.

STOA is used for instability rutting research since instability rutting is a phenomenon that
often affects pavements in their first-year of service.









Coring


The cylindrical specimens were then cored to produce the hollow cylinder specimens as

described below:

* The sawed cylindrical specimens were then cored down the middle using a rotary drill
(Figure 5-12)


Figure 5-12. Coring process to obtain hollow cylinder specimen.


* The Bulk Specific Gravity of the remaining core, is then calculated using ASTM D1 189
and D2726. The percent air void of the hollow can then be calculated from the GMM.

* With the air void content of the original cylindrical specimen and the core known, a
reasonable estimate of the air void content of the ring can be calculated using volumetric
averaging.

The preceding procedures were repeated whenever a hollow cylinder specimen was

produced.









Hollow Cylinder Testing Device

A GCTS hydraulic system and load frame is one mechanical apparatus currently

employed at the University of Florida. The GCTS hydraulic system has the ability to apply both

axial force and torque to material specimens soil, rock, concrete, or asphalt. The maximum

axial force is 5 kips and 500 in-lbs torque. The torque is applied through a horizontal hydraulic

actuator and is controlled either a load cell (torque mode) or LVDT (displacement mode).

A piston-cylinder constructed at the University of Florida, provides confinement to the

sample chamber and is controlled by the use of calibrated pressure transducers. The piston-

cylinder device has been integrated into the Testar IIm software. (The piston-cylinder uses water

as the fluid medium.) Thus, axial forces, torsional forces, and confinements can be applied to

material specimens simultaneously under pre-defined failure or time limits. The GCTS system is

controlled through the MTS Testar IIm software.

Data acquisition is conducted through the Testar IIm software as well. The user has the

ability to input any test procedure with the appropriate data acquisition for a given time, number

of cycles, or a determined response is reached. The GCTS hydraulic system has been used

extensively at the University of Florida for research in asphalt materials and a more detailed

description of the GCTS system and load frame can be found there [Frank 1999, Swan 1999, and

Pham 2000].

Prior asphalt concrete testing at the University of Florida consisted of solid 100mm

diameter specimens 150mm in height. Previous asphalt concrete material testing at the

University of Florida with the GCTS used 100 mm diameter and 150mm high solid specimens.

These specimens were tested in a chamber connected to the GCTS load frame and controlled by

the Testar IIm software via a personal computer. Minor modifications were necessary to

incorporate the GCTC system for hollow cylinder testing. The same Plexiglas chamber that had









been previously used for 100mm specimen testing was employed for the hollow specimen

testing.

The components of HCTD, proceeding from bottom to top of the physical device, are

outlined below with a figure below each section:

1. PEDESTAL. A hollow stainless steel Pedestal provides the foundation for the
HCTD. The Pedestal is bolted to the chamber base through two (2) pairs of
anchor bolts on flanges at the base of the Pedestal. The Pedestal has an inner
diameter of 35mm. One (1) 5mm diameter opening was threaded through the
Pedestal to allow the pressure on the outside and inside of the specimen remains
the same. (This 5mm hole is threaded and sized for a screw-plug, such that the
screw plug can be inserted if a pressure differential up to 20 psi is desired.) The
Pedestal contains four (4) equally spaced threaded bolt-holes at the top to secure
the bottom End Cap.


Figure 5-13. Plan view of the pedestal.































Figure 5-14. Side view of pedestal.

2. END CAPS


a. Bottom End Cap. A hollow stainless steel Bottom End Cap provides the
connection from the asphalt specimen to the Pedestal. The Bottom End
Cap is two-tiered. The lower tier has a 6.9-inch outer diameter and the
upper tier has an outer diameter of 59 inches. The lower tier forms a
flange for the bolts to connect to the Pedestal. There is an O-ring groove
on the End Cap side in contact with the Pedestal. The 5.9-inch diameter
side of the Bottom End Cap is in contact with the AC specimen. This side
is serrated at 20 lines/inch to provide additional frictional with the asphalt
specimen. The asphalt specimen is glued to the serrated end through a
high strength Epoxy.





























gure 3-13. lFan view or tne bottom ena cap.


figure 3-10. clae view or tne Dottom ena cap.


b. Top End Cap. A hollow stainless steel Top End Cap provides the
connection from the asphalt specimen to the Connection Cap (described
below). The Top End Cap has a 5.9-inch outer diameter w. The side of the
End Cap in contact with the asphalt specimen is serrated at 20 lines/inch to
provide additional frictional with the asphalt specimen. The asphalt
specimen is glued to the serrated end through a high strength Epoxy. The
side in contact with the Connection Cap, non-serrated side, contains an O-









ring groove. The non-serrated side also contains four (4) equally spaced
threaded bolt holes. One (1) 5mm diameter opening was threaded
through the Top End Cap to allow the pressure on the outside and inside of
the specimen remains the same. (This 5mm hole is threaded and sized for
a screw-plug, such that the screw plug can be inserted if a pressure
differential is desired as with the Pedestal.)


Figure 5-1 /. Plan view ot the top end cap.


Figure 5-18. Side view of the top end cap.











3. Connection Cap. The stainless steel Connection Cap serves as the link between
the Top End Cap, and the Loading Arm of the GCTS system the load arm
provides axial force and torque to the asphalt specimen. There are four (4) equally
spaced smooth-bored bolt guide holes on the bottom of the Connection Cap that
align with the four (4) threaded bolt holes on the Top End Cap. The top of the
Connection Cap also contains a 13mm diameter groove, 10mm in depth acting as
a guide hole for the GCTS Loading Arm. There are four (4) threaded bolt holes
surrounding the groove. These threaded bolt holes connect the Connection Cap to
the Stabilizer Collar. (The Stabilizer Collar was designed previously by others as
a stabilizer and locking device for the GCTS Load Arm. The Collar was not
modified for the HCTD but was utilized.)


Figure 5-19. Plan view of the connection cap.


Figure 5-20. Side view of the connection cap.









Water is used as the medium for pressure control and, with an attached temperature

control device, can be heated or cooled to allow testing under various temperature ranges. Latex

rubber membranes (on the outside and inside the hollow cut-out) shield the asphalt specimen

from the water. The inner membrane is held in place by the O-rings fitted between the Bottom

End Cap and the Pedestal and the Top End Cap and the Connection Cap. A seal is ensured by the

force exerted by the bolted connections. Epoxy is used to bond the asphalt specimen to the

serrated End Caps through a method that has been successfully demonstrated previous solid

torsional testing of asphalt concrete [Pham, 2003.]






Connection
Top End Cap
Cap


Bottom
End Cap



Pedestal

Threaded
Opening


Figure 5-21. The hollow cylinder testing device.

















Stabilizer
Collar





















Figure 5-22. Hollow cylinder testing device inside plexiglass chamber and surrounded by latex
membrane.

Summary

The analyses presented in Chapter 4 found that significant shear stresses at low

confinements are produced by radial tires in areas of the asphalt pavement where instability

rutting is observed to occur. The orientation of the principal planes induced by tire loads in these

areas is not aligned with the vertical and horizontal planes, but orientated 20-30 from the

horizontal. The HCTD was created to tests laboratory prepared asphalt concrete specimens under

these stress conditions. The next chapter details a testing program begun with the HCTD.









CHAPTER 6
MATERIAL SELECTION AND TESTING PROGRAM

Introduction

The research approach taken thus far has been directed at the creation of a laboratory

testing device that can more accurately replicate the stress states in the asphalt pavement that are

believed to result in instability rutting. These concepts lead to the creation of the hollow cylinder

testing device (HCTD).

Material Selection

To assess the ability of the HCTD to determine an asphalt mixture's susceptibility to

instability rutting, two (2) asphalt mixtures with known susceptibility to instability rutting were

selected for testing. Testing would be conducted on short-termed oven aged samples and at

elevated temperatures (400C) to mimic the field conditions under which instability rutting is

known to occur and propagate (new pavements during first hot season).

One mixture was selected that performed well displayed little to no instability rutting -

and one mixture was selected that performed poorly displayed signs of instability rutting. Both

mixtures were based on current Superpave mix design standards. The following table outlines the

mixes selected for the testing program and general features of each mixture:











Table 6-1. Superpave mixtures employed for testing program.
Rutting Unified Soil
Mixture Performance/Estimation Superpave Aggregate Classification
Mix Cu & Cc
Name D Type System (USCS)
Field APA IDT DesignationofAggregate
White SW
Rock SP-12.5mm
Coae 1 N/A Good Good (C ) Limestone 15.8 & 3.0 (3.6% Passing
Coarse 1 (Coarse) 20
(WRC 1) #
Superpave SW-SM
Mix #1 Poor Poor N/A (Fi) Granite 23.9 & 2.4 (5.8% passing
(SP1) )#200)


The asphalt content of the WRC 1 mix is 6.5% and the asphalt content of the SP 1 mix is

5.5% -- limestone is more absorptive than granite. PG 67-22 was used for all mixtures in this

study. The Job Mix Formula (JMF) and grain-size distribution curves of the mixes can be found

in the Appendix.

The USCS designation was placed in the above table based on the assumption that at

elevated temperatures, as the binder becomes less viscous; the aggregate frictional characteristics

of the asphalt mixture play a larger role in shear resistance. The USCS was one method of

presenting the aggregate characteristics of the aggregate portion of the mix. WRC1 is defined

as"well graded sand"; SP1 is defined as"well-graded sand with silt" based on the USGS. The

USCS designation should be viewed in conjunction with the Superpave mix designations in

terms of aggregate size and distribution. Resistance to shear is hypothesized to be critical to a

mixture's ability to resist instability rutting. (Both aggregates would be classified as A-3 sand

based on the AASHTO soil classification system which is the standard system for roadway

engineering.)









Testing Program

The purpose of the testing program for the HCTD was two-fold. The primary objective has

been stated to qualitatively and quantitatively assess an asphalt mixture's susceptibility to

instability rutting. The selection two mixes with different performances would provide a "self-

check" of the results. The second objective was to ensure the ability of the HCTD to provide

results that could be considered reliable. The ability of the HCTD to provide reliable results

would investigate the affects of coring and spatial aspects of a hollow cylinder on stress-strain.

The following tests and their objectives with the mixtures stated above were instituted:

Table 6-2. Tests performed with the hollow cylinder test device.
Test Objective Specimen Types Used in
Testing

Affects of coring and spatial Hollow Cylinder1
Axial and Torsional aspects of hollow cylinder 100mm X 150mm Solid
specimens versus solid Cylinder
Complex Modulus (G*) 2
Complex Modulus (G*) specimens at micro-strain Core
levels


Affects of coring and spatial Hollow Cylinder1
aspects of hollow cylinder 100mm X 150mm" Solid
Extension specimens versus solid Cylinder
specimens at small-strain Core2
levels

Qualitatively and
Cyclic Torsional- Quantitatively assess a Hollow Cylinder1
Extension mixture's susceptibility to
instability rutting
1 150mm Outer diameter, 70mm Inner diameter and 135 mm in height
2 -68 mm diameter and 135 mm in height solid specimen the core remaining from the
creation of the hollow cylinder









The Chapter 7 describes the Axial and Torsional Complex Modulus (E* and G*) testing

and the extension testing in more detail and presents the results of these tests; Chapter 8 details

the Cyclic Torsional-Extension testing program and results.









CHAPTER 7
COMPLEX MODULUS AND EXTENSION TESTING

Complex Modulus Testing

Background

The complex or dynamic modulus test is a test that is being investigated as a test to

predict the rutting susceptibility of hot mix asphalt mixtures by associating the asphalt mixture's

elastic and viscous response at small strains to actual field rutting potential/susceptibility.

Complex modulus testing (axial complex modulus testing) was first described as a test on hot

mix asphalt in 1962 [Pupation 1962]. Recently, NCHRP Project 9-19 evaluated the complex

modulus test and the AASHTO 2002 Design Project focused complex modulus testing as a

Simple Performance Test (SPT) for the rutting resistance of HMA mixtures [Pellian and Witczak

2002]. The dynamic modulus test is outlined in ASTM D3497.

In axial complex modulus testing, sinusoidal stress or strain amplitudes are applied

axially to an unconfined cylindrical specimen at 16, 4, and 1 HZ. The ASTM Standard also

recommends testing at temperatures of 5, 25, and 400C. The complex modulus test is similar to

the unconfined creep testing in testing set-up, but the load application and analysis of the

response differs.

The dynamic or complex modulus test relates the cyclic strain to the cyclic stress in a

sinusoidal load test. The complex modulus is defined as:


E*= o


where: oo is the stress amplitude,

So is the strain amplitude.










The complex modulus, E*, is composed of a real component known as the storage

modulus, E', and an imaginary component known as the loss modulus, E". The storage modulus

represents the elastic portion of the response and loss modulus represents the viscous portion of

the response. The storage and the loss modulus can be obtained by measuring the lag in the

response between the applied stress and the measured strain. This lag in the response is known as

the phase angle (6).

Torsional complex modulus testing is similar to axial complex modulus testing (ASTM

D3497) but instead of applying a sinusoidal stress or strain amplitude in the axial direction, a

torsional stress or strain is applied. The torsional complex modulus test at the University of

Florida applies a cyclic torsional force to the top of specimen and measures the displacement on

the outside diameter as shown in Figure 7-1. Knowing the torsional stress and strain, the shear

modulus is then calculated based on the theory of elasticity.


Torque at peak



1;1
SRotaion



HMAw c c m
Specimen
Y = Single Amplitude
Shearing Strain







Rigidly Fixed
at Bottom


Figure 7-1. Torsional shear test for hollow cylinder column.










The dynamic shear modulus is calculated from the following relationship:


G* =


where T is the applied torsional shear stress and y is the measures shear strain response.

Assuming that pure torque, T, is applied to the top of a HMA column, the shearing stress varies

linearly across the radius of the specimen. The average torsional shear stress, on a cross section

of a specimen Tavg is defined as:

Tavg = S/A

where: A is the net area of the cross section of the specimen, i.e A = 7(ro2-ri2),

ro and ri are the outside and inside radius of a hollow specimen, respectively. (For a solid

specimen, ri = 0), and

S is the total magnitude of shearing stress.

S can be calculated as:

r0
S = (2,r)dr
r,

where: Tr is the shear stress at the distance r from the axis of the specimen, i.eCr = mor/r,

where Tm is the maximum shearing stress at r = 0.

On the other hand, the torque, T, can be calculated from:

ro
T= r (2-zr)rdr = T
r
r.

where: J is the area polar of inertia, J = 7t(ro4 ri4)/2.

From above, Tm can be expressed as:

Tm = Tro/J










From the above equations, one can write the equation for zavg as:

3 3
2ro -r, T
Tavg 2 2
3r -r J

or

T
r r-
avg eq


where: req is defined as the equivalent radius. It can be seen in above equation that

req = 2/3ro for a solid specimen. req = 2/3 (ro3 ri3)/(r2 ri2) for hollow specimen. In practice, req

is defined as the average of the inside and outside radii.

Shear strain is calculated:

r eq 0
y-T
1

where: 1 is the length of specimen, and 0 is the angle of twist. The angle of twist, 0, can be

measured either using an LVDT or a proximitor,

The torsional complex modulus, G*, is composed of a real component known as the

storage modulus, G', and an imaginary component known as the loss modulus, G". The storage

modulus represents the elastic portion of the response and loss modulus represents the viscous

portion of the response. The storage and the loss modulus can be obtained by measuring the lag

in the response between the applied stress and the measured strain. This lag in the response is

known as the phase angle (6).

In addition, Poisson's ratio can be estimated when conducting axial complex modus

testing and/or axial and torsional complex modulus testing. When conducting an axial complex

modulus test lateral strains can be also be measured. With lateral and axial strains measured, the









Poisson's ratio can be determined. The shear complex modulus, G*, can be determined by the

following equation [Harvey et al. 2001]:

E*
G* =
(1 + 2v)

This equation assumes that Poisson's ratio is constant, although Poisson's ratio has been

seen as being frequency dependent [Sousa and Monismith 1987] and that linear elastic relations

with moduli hold for visco-elastic complex moduli. Recent findings using the torsional complex

modulus testing suggest that reasonable values of Poisson's ratio can be determined using

equation 7.10 from axial and torsional complex modulus testing -- E* and G* [Pham 2003].

Testing Program

Axial and torsional complex modulus tests were performed on the two (2) asphalt mixtures

described in Chapter 6 SP1 and WRC 1. Axial and torsional complex modulus tests were

performed on both hollow and solid specimens. The hollow specimens were created in the

manner described in Chapter 5. The solid specimens consisted of 100mm by 150mm specimens

created in the ServoPac gyratory compactor and from the cores from the creation of the hollow

cylinder. The axial and torsional complex modulus (E* and G*) tests were conducted using the

GCTS load frame and at a temperature of 400C. The temperature control device uses water as the

fluid medium for heating, thus all testing was performed under water with the specimens

protected from the water by latex membranes. The three specimen geometries were prepared

with the Superpave Gyratory compactor to air void content of 70.5%. The hollow cylinders air

void content was calculated from volumetric averaging of the original 150mm diameter solid

specimen and the leftover core specimen.

The torsional displacement at the top of the sample was measured by fixing a two (2) small

winged plates perpendicular to the Connection Cap- refer to Chapter 5 with the plates 1800









from each other. Two (2) LVDTs were mounted to a fixed body -- the support columns of the

chamber and would measure the displacement of the winged plates and thus the specimen. The

Testar IIm software collects the displacements from the two (2) LVDTs once testing has begun

at a sampling rate set by the user. With the displacement measured, the known dimensions of the

specimen and the torque being applied, the stress and strain can be calculated providing G*. For

axial complex modulus testing, the winged plates were removed, and the LVDTs directly

measured the axial displacement of the Top End Cap.

Data reduction was performed by methods previously generated [Swan 2000 and Pham

2003]. Depending on the sample geometry (hollow, solid core, or the 100mm solid), the axial

force and torque magnitude would vary to ensure a linear range was obtained for each specimen

geometry. Due to experimental difficulties torsional complex modulus testing was not

conducted on the solid core specimens for the WC 1 and SP1 mixes.

E* and G* Testing Results

An analysis of the results was based on the assumption that E* and G* is a material

property unique to each mixture at a given density. Thus, the means and methods in which a

specimen is made or its geometry should not influence its E* or G* value for a similar density

and any differences noticed between specimens may be assumed to be from specimen creation

techniques. Results were analyzed qualitatively with emphasis of the response between the

different methods and mixes and with little emphasis on the quantitative values. The following

table presents the air void content for each of the three specimen geometries for the two

mixtures. As previously mentioned, the air void content of the hollow cylinder was calculated by

volumetric averaging of the original 150mm diameter sample (pre-coring) and the -68mm

diameter core (post-coring).









Table 7-1. Summary of air void contents of different specimen geometries.
AIR VOID CONTENT
Specimen Geometry
SP1 WRC1

Solid 100mm Diameter 7.344% 7.359%

Hollow 7.294% 7.262%

Solid Core (-68mm Diameter) 6.641% 6.592%


0 200 400 600 800 1000 1200 1400 1600 1800 2000
Microstrain (ps)
Figure 7-2. E* versus microstrain for different WRC 1 specimen geometries under axial loading
at 1HZ.
































0 100 200 300 400 500 600 700 800 900 1000
Microstrain (py)
Figure 7-3. G* versus microstrain for different WRC 1 specimen geometries under torsional
loading at 1HZ.


1000

900


0 500 1000 1500 2000 2500
Microstrain (ps)

Figure 7-4. E* versus microstrain for different SP1 specimen geometries under axial loading at
1HZ.












200

180 -e--SP1 Hollow

160 --SP1 Solid
140 --o-SP1 Core



80


60

40

20

0
4 0 ----------------------------------


0 100 200 300 400 500 600 700 800 900 1000
Microstrain (17y)
Figure 7-5. G* versus microstrain for different SP1 specimen geometries under torsional loading
at 1HZ.


In general, the hollow cylinder and solid specimens had similar E* and G* values at higher

microstrain levels and once the E*.G* to microstrain curves became more linear. The cores, as

should be expected due to their denser state, had higher G* values that the hollow and solid

specimens. The differences in complex modulus in specimen creation is much more pronounced

at lower microstrain levels. More testing should be conducted to determine the influence of

specimen creation on complex modulus response. However, since testing in this research for

rutting potential will be conducted at higher strain levels, the observed similar response between

the different specimen creation methods allows a level of confidence that testing at higher strain

level does not warrant concern by the means and methods used to create the hollow cylinder

specimen.

A comparison between the hollow cylinder, solid, and core specimens from each mixture

to each other is presented below in the following figures.





























0 500 1000 1500 2000 2500
Microstrain (zs)
Figure 7-6. WRC1 and SP1 hollow cylinder specimen E* results versus microstrain.


0 100 200 300 400 500 600 700 800 900 1000
Microstrain (piy)
Figure 7-7. WRC1 and SP1 specimen G* results versus microstrain.

Evaluations

As mentioned before these tests were designed for a qualitative evaluation and this was

carried over when comparing the two different mixtures to each other. As indicated above, the

SP1 mixture typically is stiffer at lower microstrain levels but decreases more rapidly such that at









higher microstrain levels the WRC1 has higher G* and E* levels. The WRC1 mixtures also tend

to have less of a dramatic decrease in stiffness as the microstrain level increases compared to the

SP1 mixture. This suggests the SP1 mixture may be more of a "brittle" mix compared to the

WRC1 mixture. Further complex modulus testing is recommend to determine the subtle

difference between these two mixtures at these low strain levels.

Extension Testing

Extension tests were also performed on a WRC1 and SP1 solid, hollow, and core

specimens as was done for the complex modulus testing. The purpose of extension testing was

also primarily designed to qualitatively evaluate the different specimen creation techniques and

compare the response of the two different mixtures.

Extension testing was performed with the use of the GCTS system at 400C with both

mixtures compacted to air void contents of 7+0.5%. The specimens were tested by increasing the

lateral pressure on them such that there would be axial displacement. The pressure was increases

at a rate of 1 psi every 30 seconds. The following p-q diagram displays the pressure that was

induced on the specimens for the extension testing:


























e










0 2 4 6 8 10 12 14 16
p (psi)

Figure 7-8. P-Q diagram of stress path used for extension testing.




The following two figures display the results of the extension testing of the different


specimen configurations for the different mixtures:



25





2
-WRC1 Core

-WRC1 Solid

WRC1 Hollow
15






os




05





0 i
0 1 2 3 4 5 6
q (psi)

Figure 7-9. Axial strain versus deviator stress, q, for different WRC 1 specimen geometries.












20




SP1 Core

SP1 Hollow
15


/ /









0 1 2 3 4 5 6
Q (psi)

Figure 7-10. Axial strain versus deviator stress, q, for different SP1 specimen geometries.

Due to equipment difficulties, extension testing was not performed of the SP1 100mm diameter

solid specimen.

Based on the results of the extension tests, the WRC1 specimens showed a very similar

response. The SP1 hollow and core specimens differed to a greater degree than among the

WRC1 specimens. In general, as expected due to lower air void, the core was the stiffest of the

specimen geometries. In general, the responses indicate that different specimen geometries made

through different means do not significantly affect the material response.

Comparing the responses of the two mixtures against one another, the WRC 1 appears to

be more resistant to axial deformation than the SP1 mixture resulting in less axial strain for a

similar degree of stress increase. This can be considered consistent with the E* testing, which

indicated the WRC1 mixture was stiffer at higher microstrain levels.









Summary and Conclusions

Both mixtures specimen geometries displayed similar responses suggesting material

properties are not dependent on the means and methods used to create testing specimens. In

addition, the WRC1 mixture was seen to be a stiffer mixture than the SP1 mixture at higher

strain levels.

The next chapter seeks to analyze and evaluate the WRC 1 and SP1 mixtures under

conditions consistent with stress states found in the Chapter 4 analyses using the Hollow

Cylinder Testing Device.









CHAPTER 8
TORSIONAL EXTENSION TESTING

Introduction

This chapter discusses the response of cyclic torsional-extension tests on two Superpave

asphalt mixtures. As previously mentioned, the two Superpave mixes were selected based on

their ability to resist instability rutting WRC1 being the observed "good" performer and SP1

being the observed "poor" performer. The Finite Element analysis of Chapter 4 indicated that

high shear at low confinement occurs in areas of the asphalt pavement where instability rutting is

observed to occur when subject to radial tire contact loading.

Thus, it was decided to conduct laboratory testing under high shear stresses at low

confinement under cyclic loading conditions to simulate moving loads at 400C to simulate the

stress states where instability rutting is observed to occur and under the most adverse

environmental conditions for instability rutting. The response of each mixture under the above

conditions would determine the ability of a laboratory test to determine a mixture's susceptibility

to instability rutting. Further analysis of the response would possibly offer insight into the

mechanisms behind instability rutting and/or possible future testing programs. The tests

conducted also investigated the "healing" potential of mixtures under the above-mentioned

conditions.

Healing is a phenomenon seen in IDT resilient testing. Asphalt mixtures may "heal" or

partially heal if loading is stopped for a period of time prior to the onset of macro-damage as if

the asphalt mixture will respond as if it had never been loaded. Healing has been incorporated

into the UF cracking model for asphalt pavements. The purpose of investigating "healing" under

tests designed to investigate instability rutting was to identify the potential of an all-

encompassing damage model that could incorporate both cracking and instability.









Laboratory Loading Conditions

The following figure displays, through the use of a Mohr-Circle, the critical stress state

imposed on the samples during the cyclic torsional testing.


5 -
Horizontal Plane
Stresses
4 -
3LE m Vertical Plane
-2 Z7 Stresses



U) 1
L)

U)




-3_ 3 1 2 psi
< -1






o, = 5.7 psi
-5
NORMAL STRESS (psi)

Figure 8-1. Stress state imposed on hollow cylinder samples during testing.

The above stress states displays a shear stress of 3 psi, a lateral confining stress of 4 psi, and an

axial confining stress of 0.5 psi on the hollow cylinder specimen.

As can be seen, the stress state applied to the specimens is significantly less than the

stresses calculated by the Finite Element analysis approximately 10 times less. The stress state

values predicted by the finite element analysis ripped the asphalt specimens apart immediately at

40C after one load application cycle.

A preliminary study found that the above stress state allowed for a linear response of the

asphalt specimen until failure had been reached and was qualitatively similar to the stress state









determined from the finite element analysis similar ratio of shear stress to lateral and axial

stress. The fact that the quantitative values from the finite element (FE) analysis resulted in

immediate failure reaffirms the conclusion of the author that the computational analysis was

more of a qualitative assessment of the stress states in the asphalt pavement, rather than actual,

and demonstrates the possible limitations of the FE modeling.

As mentioned previously, tests were performed cyclically -- haversine loading at 0.5 HZ -

with the stress state in Figure 8-1 representing the maximum stress state. Loading was increased

from a hydrostatic stress state of 0.5 psi to the stress state displayed in Figure 8-1 in and then

decreased to the hydrostatic stress state in a linear manner such that the Mohr-Circle grew at a

uniform rate. No rest period was employed between each loading cycle. The following figure, in

p-q space, represents the cyclic loading conditions imposed on the hollow cylinder asphalt

specimens.

P (psi)
0 0.5 1 1.5 2 2.5
0



-1



S -2-


-3



-4

Figure 8-2. Cyclic stress path imposed on asphalt specimens for testing.









Laboratory Testing Program

A series of two tests were designed. The first series of tests was to observe the overall

response of the two mixtures under the torsional cyclic loading conditions and measure their

overall response to each other. Second, each mixture would be tested to determine if the healing

phenomenon occurs under cyclic loading conditions. Also, the magnitude of recovered strain was

investigated during the rest period to determine the portion of elastic and plastic strain. The

following table describes the tests performed on each of the mixtures under torsional cyclic

loading:

Table 8-1. Testing matrix for cyclic torsional testing.
TEST CONDITION WRC1 SP1
SPECIMENS TESTED SPECIMENS TESTED
Load to failure 2 2
Load to 90% failure, rest for
1 1
8 hours, re-load
Load to 50% failure, rest for
1 1
8 hours, re-load


Failure was defined as the state when rapid strain occurs resulting in the specimen being

torn apart typically 5 to 7% radial strain and 3.5% axial strain. After radial/axial strains of this

magnitude were reached, the specimen sheared and data collection was ceased.

Laboratory Testing Results

Testing to Failure

The first series of tests were conducted on two (2) hollow cylinder samples of the WRC1

and SP1 mixes. Both samples were short-termed aged, had air voids of 7.00.5%, and were

115mm in height. The specimens were tested at 400C.

Axial and shear movements of the top of the specimen were measured at intervals of 0.25

seconds. The axial and shear strains were then determined. In addition, the angle of axial to

radial strain (Y), as an indirect measurement of the dilation/vertical expansion of the specimen,











was also calculated; the larger the angle T, the greater the portion of axial strain to shear strain.


This angle was considered insightful based on the response of the pavements in the HVS testing


program of Chapter 3.


The following figures display the radial and axial strain response of the two mixes and the


calculated Y of the mixtures two specimens of each mix were tested, thus the fours lines on the


graphs.


10

9 .

8 SP1A
SP1 B
7 ---WRC1_A o
WRC1 B















0 200 400 600 800 1000 1200 1400 1600 1800 2000
-r











CYCLES

Figure 8-3. Shear strain versus number of cycles of two SP1 and WRC1 specimens under axial-
torsional-extension testing.


















66
4 l I


- e- SP1_A
- e- SP1_B
- WRC1_A
- -WRC1_B


29


/1


0 200 400 600 800 1000
CYCLES


1200 1400 1600 1800 2000


Figure 8-4. Axial strain versus number of cycles of two SP1 and WRC1 specimens under axial-
torsional-extension testing.



60

55

50

45
e-SP1_A
S40 e-SP1B
35 WRC1_A B
S-u-WRC1 B

S30

S25-

20

15

10

5

0


0 200 400 600 800 1000
CYCLES


1200 1400 1600 1800 2000


Figure 8-5. Angle of axial to shear strain,Y of two SP1 and WRC1 specimens under axial-
torsional-extension testing.


I
O
I
I
I


'00,0011"',JOOOOOP_










Both mixtures display a similar overall response in terms. This overall response can be

divided into three ranges: Primary, Secondary, and Tertiary. The first range, or Primary Range,

is the initial response and consists of large amounts of strain over approximately 50 to 100 load

cycles due to the initial loading. After the Primary Range, the strain rate response (slope of

curve) decreases until it reaches a minimum rate, around 200 load cycles. It now has entered the

Secondary Range. The Secondary Range is marked by a stable strain rate. Finally, the strain rate

begins to increase entering the Tertiary Range. The Tertiary Range is marked by increasing

strain rate until failure/rupture is achieved.

The following table gives the relative load cycles that the three ranges for the two

mixtures based on axial, shear, and strain response.

Table 8-2. Three ranges of the axial and shear strain response of the mixtures.
WRC1 SP1
Range (Loading Cycles) (Loading Cycles)
Primary 0 to -200 0 to -200


Secondary -200 to -1200 -200 to -1000



Tertiary -1200 to 1400(1) -1000 to 1100(1)

() Cycles to Rupture


When observing the radial strain response, both mixtures appear to be very similar,

progressing at approximately the same radial strain rate. It is only when the SP1 mix breaks off

earlier into its tertiary response does one observe the difference between the mixtures. The SP1

moves into its radial tertiary response at approximately 5% radial strain (950 to 1,025 cycles)

while the WRC1 mixtures move into their tertiary range at approximately 6.5% shear strain

(1,200 cycles).









However, in the axial strain response, the difference between the mixtures is more apparent

- the SP1 mixture has a higher axial strain rate than the WRC1 mixture. It is when both mixtures

reach approximately 2.5% axial strain do they move into their respective tertiary responses 900

to 1,000 cycles for the SP1 mixture and 1,200 cycles for the WRC1 mixture. The angle of axial

to shear strain, ', versus number of cycles in Figure 8.5 demonstrates that the SP1 is undergoing

more axial strain compared to shear strain than the WRC 1 mixture.

Load Relaxation at Tertiary Range, then Load Re-Application

The next series of tests sought to investigate the phenomenon of "healing". Loading

conditions would be the same as before. The first set of tests focused on stopping the test when

the SP1 and WRC1 specimens entered the Tertiary range --- 947 and 1,294 load cycles,

respectively -- then re-loading the specimens after a rest period of 8 hours. The stress condition

during the 8-hour rest period would be an isotropic state of stress of 0.5 psi. The following

figures present the results of these series of tests displaying axial and shear strains, and '. The

re-loading was considered to be a continuation of the test beginning at the cycle which the initial

test stopped. The strain presented for the 'reload' tests is therefore a cumulative strain.












































0 200 400 600 800 1000
CYCLES


1200 1400 1600 1800 2000


Figure 8-6. Shear strain versus cumulative number of cycles for WRC1 and SP1 specimens.



5

4.5

4 --o--SP1_C
-m-WRC1_C
3.5 SP1_Reload
3.5
---WRC1 Reload

3
z
'- 2.5
I-

J 2
X Beginning of Re-Load
1.5- Application After Rest
Period

1

0.5

0
0 200 400 600 800 1000 1200 1400 1600 1800 2000
CYCLES

Figure 8-7. Axial strain versus cumulative number of cycles.


o-SP1 C
- WRC1_C
o SP1 Reload
a--~WRCl Reload ap







Beginning of Re-Load
Application After Rest
Period















55

50
o--SP1_C
45 -- -- WRC1_C
o- SP_1Reload
S40 W--WRC1 Reload

35

S30

S25
-J



15- Application After Rest
Period
10

5

0
0 200 400 600 800 1000 1200 1400 1600 1800 2000
CYCLES

Figure 8-8. Angle 'T versus cumulative number of cycles.


The following figures display the axial and shear strain response during the loading and


during the 8-hour rest versus time on logrithimic scale.


1 10


Figure 8-9. Shear strain versus time.


100 1000 10000
Time (seconds)


100000















4.5
-- WRC1
4
SP1
3.5

3

S2.5
23% Drop
29% Drop
S2

1.5



0.5

0
1 10 100 1000 10000 100000
Time (seconds)

Figure 8-10. Axial stain versus time.



The first observation prior to resting is that both the WRC1 and SP1 specimens behaved in


a manner similar to the initial uninterrupted tests. At the end of the 8-hour rest period, both


specimens appear to recover approximately 23 to 28% of the strain. Most of this strain is


recovered after 15 minutes after the load was removed.


Upon re-loading after the 8-hour rest period both the WRC 1 and SP1 entered almost


immediately into a failure/rupture response, lasting 20 to 50 cycles despite the recovery in strain.


In the previous tests, the mixtures lasted 2 to 4 times longer in there respective tertiary ranges


prior to failure. The almost immediate jump into failure upon load re-application is similar to the


strain rates in the primary range.


Load Relaxation During Secondary Phase and Re-Loading

The third set of testing sought to stop the loading when the specimens were approximately


mid-way through their respective Secondary Ranges 503 cycles for the SP1 mixture and 650












cycles for the WRC1 mixture. Again, an 8-hour rest period was allowed prior to re-loading. Both


mixtures regained/recovered approximately 6% of their radial and axial strain during the 8-hour


rest period -- less than one-quarter of the strain recovered when the tests was stopped in the


tertiary range. The results are presented in the following figures.


16


14


12
--o--SP1
m WRC1 9
S10 -- o--SP1_Reload
z -u-WRC1_Reload

8
,)


UZ 6


4


2


0*
0 200 400 600 800 1000 1200 1400
CYCLES

Figure 8-11. Shear strain versus cumulative number of cycles.


1600 1800 2000














































0 200 400 600 800 1000 1200 140
CYCLES

Figure 8-12. Axial strain versus cumulative number of cycles.


0 200 400 600 800 1000 1200 1
CYCLES

Figure 8-13. Angle versus cumulative number of cycles.


0 1600 1800 2000


400 1600 1800 20(


0 SP1
-w WRC1
0 SP1-Reload
-*-WRCl-Reload


00


-0 SP1

---SPl- load
-WR~lReloa










Unlike the re-load application after resting during the tertiary range, the mixtures

responded as if they had not been loaded before. The following figures display the results as if

the re-loading was a separate test/specimen altogether beginning:


0 200 400 600 800 1000 1200
CYCLES


Figure 8-14. Shear strain versus number of cycles.


1400 1600 1800 2000
















4.5


4


3.5 SP1_Reload
s----WRC1_Reload
3
z
2.5
_-
2


1.5




0.5


0
0 200 400 600 800 1000 1200 1400 1600 1800 2000
CYCLES

Figure 8-15. Axial strain versus number of cycles.


60

55

50

45

40 -- SP1 Reload
4_ ---WRC1_Reload
S35

4 30

LL 25
O
20

z
15

10

5



0 200 400 600 800 1000 1200 1400 1600 1800 2000
CYCLES

Figure 8-16. Angle versus number of cycles.






159









Both mixtures responded initially in a similar manner prior to load relaxation as previously

for the separate mixtures. After the load was re-applied after the 8-hour rest period, both

mixtures responded in a very similar manner in terms of both shear and axial strain rate.

The SP1 mixture had been stopped at 500 cycles. It then was re-loaded and failed after an

additional 1,050 cycles (1,550 cumulative cycles). The WRC1 mixture was stopped at 650 cycles

and failed at 1,050 cycles (1,700 cumulative cycles). Both mixtures enter their respective tertiary

ranges at the same radial and axial strains (for the re-load test, not cumulative) as when there was

no load stoppage.

Evaluations

The hypothesis of this research centered on testing mixtures at high shear under low

confinements -- the stress state believed to result in instability rutting. In addition, the elevated

temperature ranges and relatively un-aged pavements that typically undergo instability rutting

suggest that resistance to instability rutting may be tied to a mixture's aggregate make-up.

The preliminary testing results indicated that the WRC 1 withstood the loading conditions

better than the SP1 mixture took more load applications to result in failure. It appeared that the

radial strain response of both mixtures was similar. However, the SP1 mixture did have more

initial shear strain during its primary response phase.

When both mixes reached between 2.5% and 3.0% axial strain they entered into their

respective tertiary ranges, eventually leading to failure. The SP1 Mixture entered this range at

approximately 900 to 1,000 cycles while the WRC1 mixture at about 1,200 cycles. This is more

apparent when observing the Y angle response over the duration of the test the SP1 mixture has

a high Y angle meaning a higher proportion of its strain was axially than the WRC1 mixture. It

was this axial strain that contributed to a specimen reaching failure.









Next, the phenomenon of "healing" was investigated through a series of two tests. Each

mixture had a similar response when re-loaded after being stopped during their respective

tertiary phases each went into immediate failure. This suggests that even with the large

recoverable strain, the specimens of both mixtures had achieved a state of irrecoverable

"damage". It is likely, that the tertiary state is one in which the specimens had achieved the

formation of an irrecoverable slip or failure plane within the aggregate orientation -- thus the

inability of both mixtures upon re-loading to absorb additional loading cycles -- therefore no

"healing" and failure upon re-loading. This further suggests that because of the large recovered

strain, that a threshold in terms of total strain may only be applicable during initial loading, and

not during re-loadings.

However, there was a little difference in the shear and axial strain response between the

SP1 and WRC1 mixtures when re-loading occurred after stopping during the secondary range.

Prior to load relaxation, the two mixtures were behaving as they had typically been both

mixtures had a similar radial strain response, the SP1 mixture has slightly higher axial strain rate.

Both tests were then stopped mid-way through their respective secondary response ranges and

allowed the rest under isotropic relaxed stated conditions for 8 hours.

When the load re-applied, the response between the two mixtures was almost identical in

terms of their axial and radial strain response. The SP1 mixture continued for an additional 1,050

cycles (a cumulative 1,550 cycles) and the WRC1 continued for an additional 1,050 cycles (a

cumulative 1,700 cycles). Both mixtures entered their respective tertiary ranges at approximately

2.5% axial strain when viewing the re-loading as a separate test, not as a cumulative test. Thus,

if the specimens had no "memory", the two mixtures performed almost the same and the

differences between them are almost impossible to discern which mixture is the better performer.









Conclusions

The hollow cylinder device is a testing apparatus that is useful for imposing a variety of

stress states upon asphalt specimens. This research centered on testing asphalt mixtures under

high shear at low confinement and at elevated temperature to simulate the critical conditions

which are believed to result in instability rutting. Two mixtures when tested under these

conditions performed in a manner consistent with their known resistance to instability rutting. A

mixtures's ability to resist axial strain under the critical laboratory imposed stress conditions

appeared to be was the determining factor in reaching failure. The tests also indicate that some

mixtures may improve their performance relative to other mixtures when allowed to heal during

the secondary response phase. Further research is necessary to evaluate and determine the nature

of a specific that allows it to resist instability rutting better than other mixtures. The hollow

cylinder devise created here at the University of Florida will help to investigate these issues due

to its ability to impose a variety of stress states on specimens not allowed by other testing

devices.









CHAPTER 9
CONCLUSION

Summary of Work

Instability rutting is a complex phenomenon seen in asphalt pavements. Instability rutting

is rutting which is confined only to the asphalt concrete layer, typically in the top 2 to 3 inches,

and generally occurs during the first hot summer of the pavement's life. Instability rutting in

asphalt pavements is primarily due to the lateral displacement of material within the asphalt

concrete layer.

Three-dimensional finite element modeling of actual tire contact stresses identified what

are believed to be the critical stress states the result in the propagation of instability rutting -

high shear at low confinement. These stress conditions are not identified by traditional

representation of radial tire loads as uniform vertical loads. Testing under high shear at low

confinement conditions would lead to a better understanding of the mechanisms behind

instability rutting. A testing device was therefore created that could impose these critical stress

states upon asphalt specimens --the hollow cylinder testing device (HCTD) at the University of

Florida.

Preliminary testing indicated that a "good" performing mixture might be the result of its

ability to resist axial strain (dilation and/or expansion) better than "poor" performing mixtures

under critical stress states.

A HCTD is a complex device that is primarily a research tool. The HCTD is not a

screening tool to predict instability rutting in mixtures that can enjoy widespread use nor is the

protocol for such a device ready base don the limited studies conducted to date. However, prior

to development of a simple screening device, an understanding of the mechanisms behind

instability rutting is necessary. This is such that any future device, protocol, or methodology that









is developed will be based on a sound mechanistic basis rather than an empirical one which may

or may not be capturing the essence of the mechanism behind instability rutting.

Future Work

The HCTD is a versatile laboratory tool. The creation of the HCTD will allow researchers

the ability to tests under a variety of stress states that cannot be obtained by many current

devices. Based on the results of the preliminary findings the following topics should be further

investigated:

* The influence on aggregate gradation on the ability of a mixture to resist instability rutting
under critical stress states imposed by the HCTD.

* The influence on gradation and aggregate arrangement during testing under critical stress
conditions to determine a failure "limit".

* The incorporation of "healing" and failure limits to evaluate a mixture's potential for
instability rutting.

* The verification of material response by "simpler" tests through confirmation by more
complex stress states in the HCTD.

* The use of the HCTD through internal pressure greater than the external pressure induce
hoop stress -- in the study of cracking.









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BIOGRAPHICAL SKETCH

Marc Edmond Novak was born in Rahway, New Jersey on November 14, 1975 to John and

Elaine Novak. He attended Jesuit High School in Tampa, FL and after graduation attended the

University of Florida.

Marc Novak graduated with his Bachelor of Science degree in Civil Engineering from the

University of Florida in 1999 and then joined the graduate program. He obtained his Master of

Engineering in 2000. Marc Novak was selected as an Alumni Research Fellow to pursue a

doctoral degree in the field of pavement engineering at the University of Florida.

In the fall of 1996, Marc was initiated into Tau Beta Pi where he met his future wife, Miss

Patricia Kwong. In February of 2003, Marc married Patricia at St Lawrence Catholic Church in

Tampa, Florida.

In 2004, he began to work in Tampa for a materials and geotechnical engineering

consulting firm. Upon completion of his doctoral program, Marc intends to continue in the

consulting field, enjoying time with his wife, and eventually pursue a career in research and

academia.





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CREATION OF A LABORATORY TESTING DEVICE TO EVALUATE INSTABILITY RUTTING IN ASPHALT PAVEMENTS By MARC E. NOVAK A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY UNIVERSITY OF FLORIDA 2007 1

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2007 Marc E. Novak 2

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To my wife Patricia 3

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ACKNOWLEDGMENTS I would like to acknowledge those individuals who were involved in the advancement of this research. First, I would like to thank my advisor and mentor Dr. Bjorn Birgisson for his time and support and the advice given to me. Acknowledgments also must go to my committee members for their time and input on numerous aspects of the research Dr. Reynaldo Roque, Dr. Michael McVay, and Dr. Joe Tedesco. I would also like to thank my outside committee member, Dr. Bhavani Sankar, for his input. Special thanks and appreciation go to the University of Florida staff. I would like to especially thank Mr. George Lopp for his time and expertise and graduate students including Aditya Ayithi, Linh Pham, Dinh Nguyen and Christos Drakos for the numerous discussions and debates on research, but more importantly, their friendship. Special thanks also to Jason Crockett and William Vash, their friendship, support, and laughter is greatly appreciated. In addition, I would like to thank my father and mother-in-law, John and Rubi Kwong, for their support and encouragement. I greatly appreciate the care and support of the St. Augustine Catholic Church community. I also appreciate the time and understanding afforded to me by Luis Mahiquez, Henri Jean, Jeanne Berg, and Larry Moore of Tierra, Inc. I would also like to thank my parents, John and Elaine Novak, for being a source of support throughout my studies. In addition, I appreciate the love and support of my grand-parents Michael and Helen Novak, Edmond Wojtowicz, Eleanor Wojtowicz Buber, and my step-grandfather, Michael Buber. Finally, and most importantly, I would like to thank my wife, Patricia; it was her love, patience, and support that helped me the most. 4

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TABLE OF CONTENTS page ACKNOWLEDGMENTS ...............................................................................................................4 LIST OF TABLES ...........................................................................................................................8 LIST OF FIGURES .........................................................................................................................9 ABSTRACT ...................................................................................................................................14 CHAPTER 1 INTRODUCTION..................................................................................................................16 Background.............................................................................................................................16 Problem Statement..................................................................................................................18 Hypothesis..............................................................................................................................18 Objectives...............................................................................................................................19 Scope.......................................................................................................................................20 Research Approach.................................................................................................................20 2 LITERATURE REVIEW.......................................................................................................21 Permanent Deformation..........................................................................................................21 Tire-Interface Stresses............................................................................................................23 Accelerated Pavement Testers................................................................................................27 Torture Test Devices...............................................................................................................29 Laboratory Tests.....................................................................................................................30 Creep Tests......................................................................................................................30 Complex Modulus Testing..............................................................................................32 Triaxial Testing...............................................................................................................34 Superpave Shear Tester...................................................................................................36 Hollow Cylinder Testing.................................................................................................38 Summary.................................................................................................................................44 3 HEAVY VEHICLE SIMULATOR TESTING ON ASPHALT PAVEMENTS...................46 Instability Rutting in Modified and Unmodified Pavements Under HVS Loading...............47 Rutting Depths.................................................................................................................47 Rut Propagation Analysis through Surface Profiling......................................................51 Volumetric Analysis of Asphalt Cores............................................................................55 One-way Versus Two-way Directional Loading....................................................................59 Quantification of Shear Stresses Within Asphalt Pavement...................................................64 Analytical Approach...............................................................................................................70 Summary and Conclusions.....................................................................................................75 5

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4 THE APPROXIMATION OF NEAR-SURFACE STRESS STATES IN ASPHALT CONCRETE THROUGH THREE-DIMENSIONAL FINITE ELEMENT ANALYSES....76 Background.............................................................................................................................76 Three-dimensional Finite Element Model..............................................................................77 Pavement Structure and Loading Conditions Analyzed.........................................................80 Pavement Structure..........................................................................................................80 Loading Conditions Analyzed.........................................................................................80 Load Application.............................................................................................................83 Axisymmetric model.......................................................................................................85 Three-Dimensional Solution Process..............................................................................87 Results of Three-Dimensional Analysis.................................................................................88 Significance of High Shear at Low Confinement...................................................................97 Summary and Conclusions.....................................................................................................99 5 CREATION OF A HOLLOW CYLINDER DEVICE.........................................................101 Introduction...........................................................................................................................101 Hollow Cylinder Specimen Dimensions..............................................................................102 Closed-form analysis of stresses across hollow cylinder wall...............................104 Finite element analysis of stress non-linearity and concentrations........................110 Creation of a Hollow Asphalt Concrete Specimen...............................................................115 Hollow Cylinder Specimen Production................................................................................116 Aggregate Preparation and Batching.............................................................................116 Mixing...........................................................................................................................117 Short Term Oven Aging (STOA) and Compacting.......................................................117 Coring............................................................................................................................118 Hollow Cylinder Testing Device..........................................................................................119 Summary...............................................................................................................................126 6 MATERIAL SELECTION AND TESTING PROGRAM...................................................127 Introduction...........................................................................................................................127 Material Selection.................................................................................................................127 Testing Program....................................................................................................................129 7 COMPLEX MODULUS AND EXTENSION TESTING....................................................131 Complex Modulus Testing...................................................................................................131 Background....................................................................................................................131 Testing Program............................................................................................................135 E* and G* Testing Results............................................................................................136 Evaluations....................................................................................................................140 Extension Testing.................................................................................................................141 Summary and Conclusions...................................................................................................144 6

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8 TORSIONAL EXTENSION TESTING...............................................................................145 Introduction...........................................................................................................................145 Laboratory Loading Conditions............................................................................................146 Laboratory Testing Program.................................................................................................148 Laboratory Testing Results...................................................................................................148 Testing to Failure...........................................................................................................148 Load Relaxation at Tertiary Range, then Load Re-Application....................................152 Load Relaxation During Secondary Phase and Re-Loading.........................................155 Evaluations...........................................................................................................................160 Conclusions...........................................................................................................................162 9 CONCLUSION.....................................................................................................................163 Summary of Work................................................................................................................163 Future Work..........................................................................................................................164 LIST OF REFERENCES.............................................................................................................165 BIOGRAPHICAL SKETCH.......................................................................................................176 7

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LIST OF TABLES Table page 3-1 Test sections in HVS testing..............................................................................................48 3-2 Summary of asphalt cores from pavement sections 5A and 1B post-testing.....................56 3-3 Summary of core data through use of phase diagrams......................................................58 3-4 Deformation of an element in an asphalt pavement near the edge of the tire in the YZ-plane (transverse plane)...............................................................................................63 3-5 Properties of pavement layers used in the finite element analysis.....................................66 3-6 Values used in Burgers model to represent asphalt concrete pavement............................74 3-7 Results of Burger model under two-way and one-way loading.........................................74 4-1 Material properties of the various layers used in the analysis...........................................80 5-1 Possible hollow cylinder specimen dimensions...............................................................104 5-2 Inner and outer pressures applied to proposed hollow cylinder geometries....................105 5-3 Beta 3 shear stress non-uniformity quantities for the various geometries.......................107 6-1 Superpave mixtures employed for testing program.........................................................128 6-2 Tests performed with the hollow cylinder test device.....................................................129 7-1 Summary of air void contents of different specimen geometries....................................137 8-1 Testing matrix for cyclic torsional testing.......................................................................148 8-2 Three ranges of the axial and shear strain response of the mixtures...............................151 8

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LIST OF FIGURES Figure page 2-1 Structural or consolidation rutting.....................................................................................22 2-2 Instability rutting................................................................................................................23 2-3 Vertical stress distribution of a radial tire..........................................................................24 2-4 Lateral stress distribution of a radial tire...........................................................................25 2-5 The Smithers Scientific Inc. tire contact stress measurement device................................25 2-6 SST test chamber...............................................................................................................37 2-7 Plan view of a hollow cylinder with outer radius, r o and inner radius, r i and subject to varying outer (P o ) and inner pressure (P i )......................................................................41 2-8 Profile view of a hollow cylinder specimen......................................................................41 3-1 The Mark IV Heavy Vehicle Simulator device.................................................................46 3-2 Rut depth progression for the various mixes.....................................................................49 3-3 Trench cut on one of the sections tested during the HVS study........................................50 3-4 Transverse cross-section of the asphalt pavement after testing.........................................50 3-5 Average transverse profile created by averaging the 58 longitudinal measurements from the laser profiler........................................................................................................51 3-6 The 59 longitudinal values averaged.................................................................................52 3-7 Average transverse profile with ratio equation..................................................................53 3-8 The ratio of area elevated to area depressed for all 50 C sections...................................54 3-9 Final transverse rut profile of sections 1B and 5A............................................................55 3-10 Phase diagram of the asphalt cores....................................................................................57 3-11 Comparison of one-way and two-way directional loading with no tire wander................60 3-12 Comparison of one-way and two-way directional loading with tire wander.....................60 3-13 Instability rutting with directional axes.............................................................................61 3-14 Two-way versus one-way directional loading...................................................................62 9

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3-15 Deformation of an element in an asphalt pavement near the edge of the tire in the XZ-plane (longitudinal plane)............................................................................................64 3-16 Area of pavement system where shear stresses were obtained..........................................67 3-17 Shear stress in YZ-plane....................................................................................................68 3-18 Shear stress in XZ-plane....................................................................................................68 3-19 Shear stress in XY-plane....................................................................................................69 3-20 Burgers model....................................................................................................................70 3-21 Longitudinal shear stress pattern in one-way loading........................................................73 3-22 Longitudinal shear stress pattern in two-way loading.......................................................73 4-1 Three-dimensional finite element mesh used in the analysis.............................................79 4-2 Plan view of the three-dimensional finite element mesh used in the analysis...................79 4-3 Variation in vertical stress across the 5-rib radial tire used in the analysis.......................81 4-4 Variation in vertical stress across 4-rib radial tire used in the analysis.............................82 4-5 Variation in horizontal stress across 5-rib radial tire used in the analysis.........................82 4-6 Variation in horizontal stress across 4-rib radial tire used in the analysis.........................83 4-7 Contact area and the radial tire nodal forces used in the pavement response analysis......85 4-8 Cross-sectional view of applied tire nodal forces used in the pavement response analysis...............................................................................................................................85 4-9 A cross-sectional view of the axisymmetric finite element mesh used for comparison purposes.............................................................................................................................86 4-10 Comparison of shear stresses under the edge of a two-dimensional axisymmetric circular uniform vertical load predicted with BISAR and ADINA...................................87 4-11 Location within asphalt pavement from finite element analysis for preliminary stress state evaluation...................................................................................................................89 4-12 Maximum shear stress in longitudinal direction along outer tire rib.................................90 4-13 Confinement stress in longitudinal direction along outer tire rib......................................91 4-14 Confinement and maximum shear stress path...................................................................91 10

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4-17 Maximum shear stress magnitude (in kPa) and direction under the 5-rib radial tire loading condition...............................................................................................................94 4-18 Maximum shear stress magnitude (in kPa) and direction under the 4-rib radial tire loading condition...............................................................................................................95 4-19 Maximum shear stress magnitude (in kPa) and direction under the uniform loading condition............................................................................................................................95 4-21 Stress states at point A for the three loading conditions analyzed.....................................96 4-22 Stress state at point B for the three loading conditions analyzed......................................97 4-23 Typical strength envelopes in hot mix asphalt and granular materials..............................98 5-1 Ratio of radial stress to average radial stress across the wall of the specimen for various geometries...........................................................................................................106 5-2 Ratio of tangential stress to average tangential stress across the wall of the specimen for various geometries......................................................................................................106 5-3 Ratio of shear stress to average shear stress across the wall of the specimen for various geometries...........................................................................................................107 5-4 Hollow cylinder specimen dimensions............................................................................109 5-5 Three-dimensional finite element model of the hollow cylinder specimen for end effects analysis.................................................................................................................110 5-6 Tangential (hoop) stress distribution across hollow cylinder wall from finite element analysis.............................................................................................................................111 5-7 Radial stress distribution across hollow cylinder wall from finite element analysis.......112 5-8 Three vertical locations where stress distributions across the wall from the finite element solution were compared to the closed-form solution.........................................113 5-9 Comparison of the tangential stress across the hollow cylinder wall from the closed-form solution and the finite element analysis with end constraints.................................113 5-10 Comparison of the radial stress across the hollow cylinder wall from the closed-form solution and the finite element analysis with end constraints..........................................114 5-11 Hollow cylinder specimen production.............................................................................116 5-12 Coring process to obtain hollow cylinder specimen........................................................118 5-13 Plan view of the pedestal.................................................................................................120 11

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5-14 Side view of pedestal.......................................................................................................121 5-15 Plan view of the bottom end cap......................................................................................122 5-16 Side view of the bottom end cap......................................................................................122 5-18 Side view of the top end cap............................................................................................123 5-19 Plan view of the connection cap......................................................................................124 5-20 Side view of the connection cap......................................................................................124 5-21 The hollow cylinder testing device..................................................................................125 5-22 Hollow cylinder testing device inside plexiglass chamber and surrounded by latex membrane.........................................................................................................................126 7-1 Torsional shear test for hollow cylinder column.............................................................132 7-2 E* versus microstrain for different WRC1 specimen geometries under axial loading at 1HZ..............................................................................................................................137 7-3 G* versus microstrain for different WRC1 specimen geometries under torsional loading at 1HZ.................................................................................................................138 7-4 E* versus microstrain for different SP1 specimen geometries under axial loading at 1HZ..................................................................................................................................138 7-5 G* versus microstrain for different SP1 specimen geometries under torsional loading at 1HZ..............................................................................................................................139 7-6 WRC1 and SP1 hollow cylinder specimen E* results versus microstrain.......................140 7-7 WRC1 and SP1 specimen G* results versus microstrain................................................140 7-8 P-Q diagram of stress path used for extension testing.....................................................142 7-9 Axial strain versus deviator stress, q, for different WRC1 specimen geometries...........142 7-10 Axial strain versus deviator stress, q, for different SP1 specimen geometries................143 8-1 Stress state imposed on hollow cylinder samples during testing.....................................146 8-2 Cyclic stress path imposed on asphalt specimens for testing..........................................147 8-3 Shear strain versus number of cycles of two SP1 and WRC1 specimens under axial-torsional-extension testing...............................................................................................149 12

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8-4 Axial strain versus number of cycles of two SP1 and WRC1 specimens under axial-torsional-extension testing...............................................................................................150 8-5 Angle of axial to shear strain, of two SP1 and WRC1 specimens under axial-torsional-extension testing...............................................................................................150 8-6 Shear strain versus cumulative number of cycles for WRC1 and SP1 specimens..........153 8-7 Axial strain versus cumulative number of cycles............................................................153 8-8 Angle versus cumulative number of cycles.................................................................154 8-9 Shear strain versus time...................................................................................................154 8-10 Axial stain versus time.....................................................................................................155 8-12 Axial strain versus cumulative number of cycles............................................................157 8-14 Shear strain versus number of cycles...............................................................................158 8-15 Axial strain versus number of cycles...............................................................................159 8-16 Angle versus number of cycles....................................................................................159 13

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Abstract of Dissertation Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Doctor of Philosophy CREATION OF A LABORATORY TESTING DEVICE TO EVALUATE INSTABILITY RUTTING IN ASPHALT PAVEMENTS By Marc E. Novak August 2007 Chair: Bjorn Birgisson Major: Civil Engineering Near-surface rutting within the asphalt layer, known as instability rutting, has become a costly mode of pavement distress on todays roads. Instability rutt ing reduces ridability, increases the potential for ponding leading to the potential for hydroplaning, and necessitates costly rehabilitation. This research sought to id entify the mechanisms behind instability rutting and to development a laboratory testing device that could evaluate an asphalt mixtures ability to resist instability rutting. Three-dimensional finite element analysis using the program ADINA was used to identify stresses radial tires induce to the pavement. The analyses indicat ed that radial tires induce high shear stress at low confinements within the asphalt pavement in areas where instability rutting is observed to occur at levels not predicted by trad itional uniform vertical loading. High shear at low confinement is believed to be a key factor in the mechanism of instability rutting. Based on the results of the finite element analys es, a laboratory device that could replicate the high shear stresses at low confinements was sought. A laboratory device that could achieve this was the hollow cylinder testing device. By testing asphalt specimens under the critical stress 14

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condition of high shear at low confinement, it is believed will help in evaluating a mixtures ability to resist instability rutting. A hollow cylinder testing device was developed at the University of Florida within the frame work of existing laboratory equipment. The hollow cylinder device was developed such that axial, shear, and confinement stresses (inner and outer pressures) could be applied simultaneously to induce stress states similar to those identified by the finite element analyses. Laboratory prepared asphalt specimens of known instability rutting performance were tested with the hollow cylinder device under cyclic stress critical stress and axial and shear strains were measured. The hollow cylinder is a device that will help lead to a possible screening tool to determine a mixtures susceptibility to instability rutting and will lead to insight into the mechanisms behind instability rutting. 15

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CHAPTER 1 INTRODUCTION Background A major distress mode in flexible (asphalt concrete) pavements is rutting. Rutting is the mechanism that produces depressions in the wheel-paths of asphalt concrete pavements. Rutting is the result of volumetric compression and/or shear deformation of one or more layers of the pavement system (asphalt concrete, base, and/or subgrade) under repeated traffic loadings. Rutting reduces serviceability and creates the potential for hydroplaning due to the accumulation of water in the wheel-path ruts. One form of rutting is known as instability rutting. Instability rutting is rutting which is confined only to the asphalt concrete layer. Instability rutting in asphalt pavements is primarily due to the lateral displacement of material within the asphalt concrete layer. Instability rutting is generally seen in pavements with a thick asphalt concrete layer (high trafficked roadways) and is the predominant mode of premature failure is modern asphalt pavements. Instability rutting is attributed strictly to the asphalt mixture properties and usually occurs within the top 2-3 inches of the asphalt concrete layer. Instability rutting occurs when the structural properties of the compacted pavement are inadequate to resist the stresses imposed upon it. Despite instability rutting being the predominant mode of premature rutting failures in modern flexible pavements, current pavement structural design approaches do not deal with rutting in the asphalt concrete layer. In 1987, Congress established the Strategic Highway Research Program (SHRP) with the objective of improving the performance and durability of roadways in the United States. The Superpave (Superior Performing Asphalt Pavements) mix design method was the result of the SHRP to create a rational mix design method that could minimize distress in asphalt concrete 16

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pavements, including rutting. The Superpave mixture design, although an improvement on previous mixture design methods, is based solely on volumetrics. In Superpave mix design, an asphalt mixtures susceptibility to instability rutting is mitigated by ensuring proper volumetric quantities of the asphalt mixture are satisfied. Superpave has yet to incorporate a test that directly measures or evaluates a mixtures resistance to rutting. The desire for a simple and direct quantitative and/or qualitative measure of a mixtures susceptibility to instability rutting is currently being sought in conjunction with the current Superpave mix design. Today, most of the intra-continental commerce in the United States is transported through trucks as opposed to rail in years past. In 1988, 2.2 million miles of our nations roadways had asphalt concrete surfaces with 91% of 2 trillion annual vehicle miles occurring on these asphalt concrete roadways [Federal Highway Administration (FHWA) 1988]. Since that time, this number has increased. Over the last 30 years, the type of tires trucks are using is changing as well. Today, more than 98% of trucks use radial tires, as opposed to bias-ply in years past. Radial tires allow for higher inflation pressures and thus, the ability to carry higher loads. Studies have shown that radial tires impart quantitatively and qualitatively different stresses to asphalt pavements [Myers 1997]. It has been shown that radial tires are more detrimental to the pavements surface than previous bias-ply tires [Myers et al. 1999]. Numerous performance prediction tests have been presented, discussed, and implemented by different transportation agencies in order to determine an asphalt mixtures susceptibility to instability rutting. In general, prediction tests may either be considered physical or index. Physical tests can be simply defined as torture tests. Torture tests subject an asphalt concrete pavement or specimen to loading conditions that mimic field conditions. Large scale torture 17

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testing is done on actual pavement sections through the use of vehicle simulators. Small scale torture testing is done with loaded wheel testers (LWTs). LWTs subject a laboratory created asphalt specimen to repeated miniature moving wheel loadings. The measured response (degree of deformation) is then linked to field performance (rutting). Some common LWTs include the Hamburg Wheel-Tracking Device, the French Pavement Rutting Tester, and the Asphalt Pavement Analyzer. Index tests are generally laboratory tests that attempt to determine one or more fundamental material properties of asphalt concrete or the response of asphalt concrete under certain loading conditions and link that property or response to overall rutting performance. Such laboratory index tests that are concurrently being employed are creep (cyclic and static), complex modulus, triaxial, and hollow cylinder tests. Unlike LWTs, laboratory tests place an equivalent stress state throughout the specimen and can measure fundamental properties of the asphalt mixture. Both approaches have merits and deficiencies and are currently being evaluated. Problem Statement Factors that influence or contribute to the amount of rutting and/or an asphalt concrete pavements resistance to rutting have not been clearly identified. Because of this lack of understanding of the factors that influence instability rutting, no effective way of predicting and evaluating the rutting potential of asphalt mixtures exists today. There is a need to identify the critical factors that may contribute to instability rutting. Once identified, current laboratory testing devices can be evaluated or new tests designed to determine a mixtures instability rutting susceptibility. Hypothesis Instability rutting is primarily the result of high near-surface shear stresses at low confinements. The high near-surface shear stresses at low confinement within the asphalt 18

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pavement driving rutting is the result of the stress distribution in todays high inflation radial tires. Instability rutting may therefore be controlled by the shear strength and stiffness of the material at low confinement. The shear strength and stiffness of the materials at low confinement are influenced by a combination of material state (density), loading history (compaction), and environmental conditions (temperature). Thus, a laboratory test that can best replicate this set of critical conditions could be used to determine key material properties for the evaluation of instability rutting in asphalt mixtures. Objectives The main objective of this research is to identify the mechanism of instability rutting at critical conditions that can lead to a laboratory instability rutting prediction procedure. The only laboratory test that can replicate high shear stress at low confinement is the hollow cylinder test. The primary objectives of this research study are listed below: Evaluation of instability rut progression under controlled field conditions Identification of material stress states under actual tire contact stresses Design and construct a new laboratory testing device (hollow cylinder test) to induce realistic stress states on laboratory asphalt concrete specimens Verify laboratory rutting response to field rutting response Identification of material response under critical loading and material conditions using the hollow cylinder device Recommend testing configurations and methods for mixture rutting evaluation. Instability rutting is a complex phenomenon in the area of asphalt pavement distress. The culmination of these objectives will eventually, after further research and experimentation, lead to the development of a screening tool or design methodology that can predict the amount of instability a mixture may undergo under typical loading and environmental conditions. The research presented here is a first attempt to begin to get our hands around the very complex 19

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phenomenon of instability rutting and help pave the way for our ability to predict and understand instability rutting. Scope The research focuses on identifying the critical conditions that contribute to the mechanism of instability rutting. Defining the conditions that initiate and propagate instability rutting will lead to the development of better prediction tests and models. It is not possible to examine all possible parameters that affect rutting of asphalt concrete mixtures in the allotted time frame. Research Approach This research is divided into two parts the analytical and the experimental. The analytical part includes a three-dimensional finite element analysis (FEA) of radial tire contact stresses and an evaluation of the critical stress states believed to be driving instability rutting. The experimental part includes the creation of a laboratory testing device, the testing, and data analysis. A research-approach outline is presented below: Literature review: examine existing ideas, concepts, theories and results published on tire contact stresses, rutting in asphalt pavements, torture test devices and laboratory test devices used in rut prediction Field evaluation: examine the rut progression of two mixtures under High Vehicle Simulator (HVS) loading conditions and under two different loading conditions to understand the propagation of instability rutting under quasi-actual conditions. Tire contact stresses: employ three-dimensional finite element modeling to estimate the near-surface stress states in asphalt concrete under radial tire contact stresses. Laboratory testing device: design and construct a new laboratory-testing device, the hollow cylinder test (HCT), to induce stresses that would be representative of the actual stresses induced by radial truck tires from the tire contact stress study. Test new device: test two mixtures with known field and laboratory performance good and poor in the HCT device and evaluate its ability to produce reliable results Evaluation and observation of tests results Discussion on the on the results obtained and recommended future testing methods 20

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CHAPTER 2 LITERATURE REVIEW This chapter reviews some of the literature available on the subjects of permanent deformation, tire-pavement interface stresses, accelerated pavement testers, torture test devices, and laboratory methods for predicting mixture performance. Permanent Deformation Rutting is one of the major modes of distress in asphalt pavements today. Rutting is defined as the longitudinal depression that forms within the wheel-path. Rutting can lead to the premature deterioration of roadways requiring costly rehabilitation and can be a significant safety concern with skidding and/or hydroplaning when water collects in the ruts. There are three types of rutting: wear, structural, and instability [Dawley et al. 1990]. The first is wear rutting, which is due to progressive loss of coated aggregate particles from the pavement surface. This is also known as stripping. The second is structural or consolidation rutting. This is the traditional term used when discussing rutting. It refers to volumetric compression and/or shear deformation of the base or subgrade with an assumption that the asphalt concrete layer contributes very little to the overall rutting of the pavement system --the bituminous layer conforms to the shape of the lower layers [Huang 1993]. This mode of rutting may result from possible insufficient compaction of base and subgrade layers, which undergo air void reduction and shear deformation under repeated traffic loadings. It can also be due to the consolidation phenomenon in clayey bases and subgrade. Structural or consolidation rutting will occur over the design lifetime of the pavement system and is not typically premature failure mode -unless the base and subgrade are poorly compacted. Rutted roads due to this mechanism (Figure2-1) are marked by shallow sloping ruts that are fairly wide (3040 inches) [Huber 1999]. 21

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Figure 2-1. Structural or consolidation rutting. (Reprinted with permission from Huber, G.A, Methods To Achieve Rut-Resistance Durable Pavements, Synthesis of Highway Practice 274, Transportation Research Board, National Research Council, Washington D.C., 1999.) The third type of rutting is instability rutting. Instability rutting is due to lateral displacement of material within the asphalt concrete layer only. Instability rutting is a near-surface phenomenon occurring in the top 2 inches of the asphalt layer [Dawley et al. 1990]. Instability rutting occurs when the structural properties of the compacted pavement are inadequate to resist the stresses from frequent repetitions of high axle loadings. The aggregates rigidly translate and rotate within the asphalt binder [Wang et al. 1999]. Instability rutting (Figure 2-2) is characterized by steep longitudinal ruts in the pavement with humps of material on either side of the rut [Huber 1999]. 22

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Figure 2-2. Instability rutting. (Reprinted with permission from Huber, G.A, Methods To Achieve Rut-Resistance Durable Pavements, Synthesis of Highway Practice 274, Transportation Research Board, National Research Council, Washington D.C., 1999.) Tire-Interface Stresses Over the past thirty years average truck tire inflation pressures have been increasing. During the AASHO road tests in the 1960s, tire inflation pressures were 75 to 80 psi, today tire inflation pressures are in excess of 100 psi [Wang et al. 2003]. The type of truck tires on the roads today is also different form twenty years ago. In the past, trucks utilized bias-ply tires, which tended to have high wall stiffness and a flexible footprint. Currently, more than 98% of trucks use radial tires, because of the associated fuel savings and higher reliability of newer tire structures [Myers et al. 1998, Roque et al. 1998]. Analyses of the tire contact stresses imparted by different tire configurations have been studied in the laboratory by many researchers. One of the earliest measurement systems employed electronic pick-ups embedded into the pavement and recorded local forces of a rolling bias-ply tire [Bode 1962]. In another study, a rotating steel drum was used to measure automobile tire contact stress [Seitz and Hussman]. It was found in both cases that tire pressures varied across the tire footprint and longitudinal stresses were present not uniform stress distribution and only vertical stresses. 23

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Researchers have also measured the normal and tangential tire-pavement interface stresses by employing a steel-bed transducer array [Woodside et al. 1992]. The Vehicle-Road Pressure Transducer Array (VSPTA) developed in South Africa measured three-dimensional tire-pavement interface stresses with 13 triaxial strain gauge steel pins (spaced 17mm transversely) mounted on a steel plate and fixed flush with the road surface [De Beer et al. 1997]. Some results from VSPTA study (Figures 2-3 and 2-4) display the non-linearity of the vertical and lateral stress distribution under a radial tire. Figure 2-3. Vertical stress distribution of a radial tire. (Reprinted with permission from De Beer M., C. Fisher, and F. Jooste, Determination of Pneumatic Tire/Pavement Interface Contact Stresses Under Moving Loads and Some Effects on Pavements with Thin Asphalt Surfacing Layers, Proceedings of the Eighth International Conference on Asphalt Pavements, Seattle, Washington, 1997, pp.179-226.) 24

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Figure 2-4. Lateral stress distribution of a radial tire. (Reprinted with permission from De Beer M., C. Fisher, and F. Jooste, Determination of Pneumatic Tire/Pavement Interface Contact Stresses Under Moving Loads and Some Effects on Pavements with Thin Asphalt Surfacing Layers, Proceedings of the Eighth International Conference on Asphalt Pavements, Seattle, Washington, 1997, pp.179-226.) Dr. Marion Pottinger of Smithers Scientific Services, Inc. developed another device (Figure 2-5) to measure tire-pavement interface stresses. 16 Transducers Bed Motion Tire RollingDirectionBed z,zy,y x,x Tire Coaxial Load and Displacement Transducer Detail 16 Transducers Bed Motion Tire RollingDirectionBed z,zy,y x,x Tire Coaxial Load and Displacement Transducer Detail Figure 2-5. The Smithers Scientific Inc. tire contact stress measurement device. 25

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By using triaxial load pin transducers inserted onto a flat steel test track, tire-interface forces and displacements for vertical, longitudinal, and transverse axes were able to be measured [Pottinger 1992]. The experimental setup used was also capable of determining the rolling tire footprint shape through the implementation of a rolling steel treadmill device in which the tire was held in one location, while the bed was moved longitudinally, causing the tire to roll over a row of 16 transducers. Stresses and displacements were recorded every 0.20 inch longitudinally (parallel to wheel path) and every 0.15 inch transversely (perpendicular to wheel path) by varying the transverse position of the sensors. The measurements provided a higher definition of actual tire-pavement contact stresses than previously obtained. The VRSPTA also measures contact stresses in the x, y, and z directions, but uses only 13 triaxial strain gauge steel pins, mounted on a steel plate and fixed flush with the road surface. Based on contact stress measurements, researchers have identified two distinct types of contact stress effects that exist under truck tires. These are generally referred to as the pneumatic effect and Poissons effect [Pottinger 1992, Myers et al. 1998]. The overriding effect induced under bias-ply tires is the pneumatic effect, and under radial truck tires is the Poissons effect. This is a direct result of tire construction. Radial tires are constructed to have stiff treads and flexible sidewalls, to minimize the deformation of the tire during rolling. Thus, the lateral stresses induced on the road by the radial truck tire will tend to push out from the center of the tire ribs. In contrast, bias-ply tires tend to have high wall stiffness and a flexible tread, resulting in smaller lateral contact stresses. Because of the complexities involved in measuring contact stresses under tires, it is not possible to obtain these measurements directly on real pavements. In particular, the question arises as to whether stresses measured under a tire on a rigid foundation with embedded sensors 26

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are similar to the contact stresses that a flexible pavement will experience. This question was answered by a two-dimensional finite element model of a radial tire [Roque et al 2000], using the finite element program ABAQUS [Hibbit et al. 1997]. The results showed that contact stresses were nearly identical whether the contact surface was rigid or had properties similar to an asphalt pavement. Effectively, the stiffness of the tire is so much less than the pavement, that the resulting contact stresses are similar to those obtained from a rigid steel bed. It was concluded that contact stress measuring devices with rigid foundations, such as Pottingers device, are suitable for the prediction of response of flexible highway pavements. Thus, radial tires impart not only vertical stress to the surface, but also significant lateral stresses. This to shift radial tires and their increasing tire pressures and the lateral stresses they induce is theorized to be a component of the greater prevalence of instability rutting in recent years [Myers et al. 1999, Drakos et al. 2001]. Accelerated Pavement Testers Accelerated pavement testing is the application of a controlled prototype load to a controlled prototype or actual pavement system. Full-scale accelerated pavement testing began with the AASHO road test in the 1960s. These tests employ actual roadways under full-scale loadings. They require a lot of space to construct such tracks or available roadways, but can analyze many different pavement systems and are realistically loaded by vehicles. This method of testing continues today with the WESTRACK test track and the NCAT test tracks. The disadvantage of this type of testing is that it also does not allow for quick evaluation of results. To limit the amount of space needed for full-scale testing and for faster analysis of results Accelerated Pavement Testers (APTs) were designed. APTs apply a simulated load to a pavement system. APTs may either be stationary or mobile devices that apply repeated wheel loadings in a linear or circular manner. Testing under 27

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these accelerated conditions allows for the quick evaluation and response of a pavement system [Metcalf 1998]. The appeal of APTs lies in their ability to produce realistic wheel loadings, including wheel wander, to a pavement system. APTs can simulate twenty years or more of traffic in a reduced period of time. There are many types of APTs in the world today. These include the Simulated Vehicle and Loading Emulator (SLAVE), the Texas Mobile Load Simulator (TxMLS), Accelerated Load Pavement Facility (ALPF) Tester, Accelerated Loading Facility (ALF) Tester, and the Heavy Vehicle Simulator (HVS). DYNATest, a South African company, developed the HVS. The HVS consists of a large carriage where a wheel rolls underneath the carriage along the pavement. The load is applied in a linear manner, and can provide lateral wheel wander and one-way or two-way directional wheel loadings. The HVS is fully mobile [Kim 2002]. An extensive program was carried out in South Africa using the HVS. The HVS was brought to different roadway sites in South Africa to test different pavement sections. The vast amount of data collected through this testing program formed the basis for the South African pavement design system [National Institute for Transport and Road Research 1985]. In order to conduct a full-scale test with any APT device, a test track must be constructed or an available roadway must be available. APTs are also expensive to manufacture (between 1 to 2 million dollars), thus are in limited quantity and require large funding. The primary disadvantage of APT is this expense and the need to construct appropriate testing facilities. Furthermore, aging effects are limited, and environmental conditions are not controllable without special facilities. Despite these limitations, the insight gained with proper measurements and analysis is worthwhile and can evaluate the effectiveness of other tests. 28

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Torture Test Devices Torture testing of asphalt specimens has been gaining wide appeal recently for their relative ease and simplicity. These tests subject an asphalt concrete laboratory prepared specimen to repeated loadings. One method of load application is through small wheel loadings rolling over an asphalt specimen. Torture tests of this method are known as Loaded Wheel Testers (LWTs). LWTs can be viewed as a small scale APTs. The primary purpose of LWTs is to perform efficient, effective, and routine laboratory rut proof testing and field production quality control of asphalt mixtures [Lai 1990]. There are many types of LWTs today. Europeans have developed the Hamburg Wheel-Tracking Device (HWTD) and the French Pavement Rutting Tester (FPRT). Americans have developed at the University of Arkansas the Evaluator of Rutting and Stripping in Asphalt (ERSA) and The Georgia Department of Transportation (GDOT) began the development of a Georgia Loaded Wheel Tester (GLWT) in 1985. The Asphalt Pavement Analyzer (APA) is a modification of the GLWT [Lai, 1986]. The concept of pavement evaluation is simple when using LWTs the rut depth from the LWT can be related to field performance. Despite the advantages of simplicity and cost effectiveness, there are some issues. LWTS can usually, but not always, differentiate between good and bad performing mixtures [Collins 1996, Stuart and Mogawer 1997]. A comparison of LWTs to APTs showed that LWTS could distinguish good and bad performing mixtures when only the binder varied. However, when aggregate gradations were varied, none of the LWTs were able to distinguish between good and bad performing mixtures [Romero and Stuart 1998]. An analysis of the HWTD and FPRT noted that the devices could discriminate between mixtures with widely different binder grades, but failed to give consistent results for mixes with closer binder grades [Stuart and Mogawer 1997, Stuart and Izzo 1995]. The APA was also sensitive to mixtures with different 29

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asphalt binder and varying gradations [Kandhal and Cooley 1999]. The APA ranked WESTRACK performance to an 89% accuracy level [Williams and Prowell 1999]. Other issues of concern include the loading mechanism (pressurized hose or rubber strip or miniature wheel) the boundary conditions (rigid confinement of the laboratory prepared asphalt concrete specimen in the testing mold), and the aspect ratio of the loading strip compared to aggregate size [Federal Highway Administration (FHWA) 1998]. Laboratory Tests Creep Tests Uniaxial creep tests are tests where a cylindrical asphalt concrete specimen is subject to a load in the axial direction with no lateral confinement. This is a single point test -a single predetermined loading condition. The load may be applied statically for a predetermined duration or until a certain deformation has been achieved. The load may also be applied cyclically in a square wave load or half-sinusoidal loading pattern with rest periods between loadings. Although traditionally creep means deformation under constant load, the term dynamic creep is used in pavement mechanics and denotes a load magnitude that is constant but is applied in intervals. Creep tests can determine the creep compliance, relaxation modulus, and creep resistance of asphalt mixtures. Analyses of creep tests provide elastic, plastic, and viscous properties of mixtures that can be associated to rutting. The uniaxial static creep test has been defended as a test that is effective in identifying the sensitivity of asphalt concrete mixtures to permanent deformation with many depth prediction methodologies based on this test [Little et al. 1993, Marks et al. 1991 and Hills et al. 1974]. The Shell Method is one of the most recognized asphalt concrete mix design methodologies that was based on the static creep test [Van de Loo 1976]. However, the Shell Method does not incorporate strain hardening with time [Ali and Tayabji 30

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1998] and may overpredict rutting in modified asphalt mixtures [Monismith and Finn 1987]. The disadvantages of the static creep test are its simplicity and defined loading condition on the specimen. In order to better simulate loading conditions found in the field, testing under repeated load or cyclic conditions was introduced. Many researchers have found good correlations between repeated creep load tests and field performances and have developed rut prediction models [El Hussein and Yue 1994, Qi and Witczak 1998]. The creep test is shown to be sensitive to mixture variables including asphalt grade, binder content, aggregate type, air void content, temperature of testing, testing stress level, and rest periods [Little et al. 1993]. Any correlations made with creep testing are very dependent upon the load and rest times [Qi and Witczak 1998]. Comparisons of the asphalt mix stiffness obtained from compressive creep tests with asphalt mix stiffness obtained from wheel-tracking tests observed good agreement between the two techniques at high values of bitumen stiffness. However, at low values of bitumen stiffness, the results from the creep tests gave higher asphalt mix stiffness than those from the wheel tracking tests [Hills et al. 1974]. Furthermore, creep testing does not allow for a unique separation of plastic strains when working within the elasto-plasto-visco material framework. Another shortcoming of the uniaxial compression tests is that no direct separation can be made in response of the material to hydrostatic and deviatoric stress [Kim et al.]. The major limitation for uniaxial creep testing, both static and cyclic, is the lack of lateral confinement. Lateral confinement has been seen to be critical for proper analysis of rutting susceptibility of open graded or mastic mixtures [Little et al. 1993]. Although simplicity is the advantage of this test, the predetermined method of loading, 31

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lack of confinement, and divergent results make this test difficult to employ as a simple reliable quantitative test for instability rutting prediction. Complex Modulus Testing The dynamic or complex modulus test is another test often used to predict the rutting susceptibility of hot mix asphalt mixtures. Complex modulus testing was first described as a test on hot mix asphalt in 1962 [Papazian 1962]. Recently, NCHRP Project 9-19 evaluated the complex modulus test and the AASHTO 2002 Design Project focused complex modulus testing as a Simple Performance Test (SPT) for the rutting resistance of HMA mixtures [Pellian and Witczak 2002]. The dynamic modulus test is outlined in ASTM D3497. In complex modulus testing, sinusoidal stress or strain amplitudes are applied axially to an unconfined cylindrical specimen at 16, 4, and 1 HZ. The ASTM Standard also recommends testing at temperatures of 5, 25, and 40C. The complex modulus test is similar to the unconfined creep testing in testing set-up, but the load application and analysis of the response differs. The dynamic or complex modulus test relates the cyclic strain to the cyclic stress in a sinusoidal load test. The complex modulus is defined as: 00* E Where: 0 is the stress amplitude, 0 is the strain amplitude. The complex modulus, E*, is composed of a real component known as the storage modulus, E, and an imaginary component known as the loss modulus, E. The storage modulus represents the elastic portion of the response and loss modulus represents the viscous portion of the response. The storage and the loss modulus can be obtained by measuring the lag in the 32

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response between the applied stress and the measured strain. This lag in the response is known as the phase angle (). When conducting an axial complex modulus test lateral strains can be also be measured. With lateral and axial strains measured, the Poissons ratio can be determined. The shear complex modulus, G*, can be determined by the following equation [Harvey et al. 2001]: 21*EG This equation assumes that Poissons ratio is constant, although Poissons ratio has been seen as being frequency dependent [Sousa and Monismith 1987] and that linear elastic relations with moduli hold for visco-elastic complex moduli. Recent findings using torsional complex modulus testing suggest that reasonable values of Poissons ratio can be determined from E* and G* [Pham 2003]. Torsional complex modulus testing provides G* directly. Torsional complex modulus applies a torsional sinusoidal stress or strain amplitudes are applied axially to an unconfined cylindrical specimen rather than an axial stress or strain amplitude. The axial complex modulus has been shown to have a good correlation with the rutting resistance of HMA paving mixtures [Witzak et al. 2002]. Numerous research groups have shown that the complex modulus test can be used to characterize the temperature dependency of mixtures and viscosity characteristics over time [Witczak et al. 2002, Perraton et al. 2001]. Analysis methods for the characterization of rutting resistance in asphalt mixtures with the complex modulus test have been presented [Majidzadeh et al. 1979; Shenoy and Romero 2002]. The complex modulus test is still only a measure of the visco-elastic properties of an asphalt mixture. Since the interpretation of the complex modulus is based on the assumption of linear viscoelasticity of the mixture, it is necessary to maintain a fairly low strain level during testing to avoid any non-linear effects. Maintaining a stress level that result in a strain response that is 33

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close to linear is critical to achieve a test that is reproducible and allow for proper analysis. Strain amplitudes of 75 to 200 microstrain are suggested in order to maintain linearity during triaxial compression testing [Witczak et al. 2000]. As a drawback, complex modulus testing does not look at large strain creep or plastic strains or particle movement that truly capture rutting in the field. Triaxial Testing Triaxial testing allows for many possible stress states on an asphalt specimen. Virtually any combination of hydrostatic and deviatoric stresses can be induced on a specimen in the triaxial test [Kim 1997]. Confinement can be varied and depending on the method of loading tests may either be in compression or extension. Loading can either be strain or time dependent, as well as static or cyclic. Early researchers found many similarities between soil and asphalt mixtures when testing under static triaxial conditions. Both soil and asphalt concrete exhibited cohesion and pressure dependency [Goetz and Schaub 1951, Nijboer 1948, Gandhi and Gallaway 1967]. These early triaixal tests were conducted by applying a constant confinement to a cylindrical asphalt specimen and then loading in the axial direction at prescribed rate. Mix design methods using the static triaxial test and modeling asphalt concrete as a Mohr-Coulomb material soon followed [Smith 1949 and McLeod 1952]. Today triaxial testing has been used to determine rutting susceptibility of mixtures [Krutz and Sebally] by comparing the failure envelopes of different mixtures and new mixture design techniques still continue using static triaxial testing [Mahboud and Little 1988]. These tests are limited to measuring modulus and strength under static compression. These tests are static tests, and are not representative of the repetitive loading experienced in the field by the asphalt concrete layer. A desire for more realistic loading led to cyclic or repeated-load triaxial testing. 34

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Repeated-load triaxial tests apply an axial stress that varies with time in manner more representative of field loading conditions. Various pulse loadings have been used depending on the available equipment, but generally a sinusoidal or triangular loading pattern is employed for its close approximation to actual field loading and to avoid sudden stress changes [Brown 1976]. In fact it has been shown that for the same degree of stress, a square loading produces greater strain than sine wave [Brown and Cooper, AAPT]. This is likely due to the sudden impact of a square wave compared to a sine waves gradual build-up. In the United States, most loading is 20 to 30 repetitions/minute, with a pulse of 0.1 seconds in duration [Brown 1976]. Investigations conducted at the University of Nottingham suggest that rest periods are not significant for permanent strain tests on asphalt mixtures. Repeated triaxial testing allows the visco-elastic properties of an asphalt material to be obtained [Rowe et al. 1995]. These parameters are directly related to rutting and enable predictions of deformation in the field and comparisons to be made between different materials. The repeated triaxial testing has been considered by some to be the best practical method for the testing of pavement materials against rutting [Brown 1976]. This is because of its ability for various stress and loading conditions that best simulate actual field conditions. Some models have been proposed for rutting using the repeated triaxial test [Meyer et al. 1976]. Despite the advantages, little repeated triaxial testing is performed to asphalt concrete [Brown 1976]. Repeated triaxial testing is primarily used for bases and sub-grades. The resilient modulus (M R ) of the subgrade is the basis of the AASHTO asphalt concrete mix design methodology. The major disadvantage of triaxial testing is that only principal stresses can be applied to a specimen -the orientation of principal stresses is always in the vertical or horizontal direction. 35

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Furthermore, the application of high shear at low confinement is not possible the state of stress hypothesized to be critical to instability rutting. Superpave Shear Tester The Superpave Shear Tester (SST) is a closed-loop feedback, servo-hydraulic system that can apply axial loads, shear loads, and confinement pressures to asphalt concrete specimens at controlled temperatures. The response of asphalt concrete to these loads can be used as inputs to performance prediction models [Shenoy et al. 2001]. The SST is a product of the Strategic Highway Research Project (SHRP) to evaluate asphalt mixture performance. The test was originally developed at the University of California at Berkeley [Sousa et al. 1991]. The SST was developed to address the following issues not accounted for in other tests: Dilation under shear loading Increase in stiffness with increase in hydrostatic pressure, Negligible volumetric creep, Residual permanent deformation on removal of load, Temperature and rate loading dependence, and Difference in response between creep and repeated loading. The SST has six main components: testing chamber, test control system, environmental system, hydraulic system, air pressurization system, and measurement transducers. The testing chamber includes a reaction frame and a shear table. The reaction frame is extremely rigid providing for accurate measurements. The shear table holds a specimen during testing and is capable of applying shear loads. The specimens normally have a diameter of 150 mm and a height of 50 mm; however, specimens with diameters and heights up 200 mm can be tested with only minor modifications to the system. Shear and axial loads can be applied sinusoidally, 36

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repetitively or in static creep mode under a confining pressure and at temperatures anywhere from 0 to 70C. Four Superpave tests are capable of being performed with the SST: Repeated shear at constant height (RSCH) Repeated shear at constant stress ratio Simple shear at constant height (SSH) Frequency sweep at constant height (FSCH) The AASHTO Provisional Standard TP7-94 contains a detailed description of these tests. Figure 2-6. SST test chamber. On-going research is attempting to evaluate the reliability of the SST in rut prediction. Results of a study conducted in North Carolina indicated that the RSCH test can clearly identify the wellperforming versus poorly performing mixes. The RSCH correctly predicted the non-susceptibility to rutting of one mix, where the French Pavement Rutting Tester and the Georgia 37

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loaded-wheel tester indicated early rutting of the pavement [Tayebali et al. 1999]. A comparative evaluation of rutting and stripping of two asphalt mixtures using the SST and the Hamburg wheel-tracking device (HWTD) has also been conducted. The results showed that both the SST and HWTD correctly predicted the superiority of the Superpave mixture to the Marshall mixture [Wang et al. 2001.]. A comparative study of the RSCH to the APA was also conducted. Correlations of various parameters (permanent deformation or strain, slopes, and intercepts from linear or power law regressions) from the RSCH had significant correlation with APA rut tests [Zhang, 2002]. However other research has indicated that the SST and the performance models developed with this test have some errors that made prediction impossible [Anderson, et al., 1999]. Analysis of the RSCH test indicated that the coefficient of variation of the permanent strain at 5, 000 load cycles for samples was between 10% and 20% with a suggestion that the RSCH be used for low volume road prediction only [Romero and Anderson 2001]. Other researchers have also indicated high variability when testing with the SST and results depended on the analysis method [Romero et al. 1998]. Further research has found that the SST could differentiate between asphalt binders but are not sensitive to changes in aggregate [FHWA 1998]. Hollow Cylinder Testing The hollow cylinder test (HCT) has been used since the 1930s to study soil specimens subject to pure shear. Since then, the HCT has become more versatile with researchers now applying axial, torsional, and internal and external pressures [Saada and Townsend 1981]. The HCT has been used most extensively in soil testing. The versatility of stress states capable with the hollow cylinder has made it ideal for the studying of the anisotropic nature of soil. The rotation of principal stresses and torsional properties of materials has been shown to have 38

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significant effects on the strength and stiffness of soil [High et al. 1983, Miura et al. 1986]. There is a plethora of articles and literature in geotechnical engineering on hollow cylinder testing, configuration, issues, and material properties [Hight et al. 1983, Alarcon et al. 1986, Dusseault 1981, Vaid et al. 1990, Richardson et al. 1996]. Only recently, American pavement researchers have begun to utilize the hollow cylinder for the analysis of asphalt concrete for the determination of tensile, creep, and dynamic properties [Sousa 1986, Crockford 1993, Buttlar et al. 1998, Alavi and Monismith 1994]. The University of California at Berkeley conducted hollow cylinder testing on asphalt concrete using specimens with an inner radius of 3.5 inches (88.9 mm) and an outer radius of 4.5 inches (114.3 mm). The specimen height was originally 18 inches (457.2 mm). Specimens were created through the use of a special fabricated molds, compaction device, and unique methods [Sousa 1986]. These same researchers decreased the height to 8 inches (203.2 mm) with specimens compacted through a rolling wheel compaction procedure outlined in SHRP A-003. Research was done using this device to investigate the time and temperature dependent properties of asphalt concrete mixes and to examine the applicability of linear viscoelasticity to asphalt concrete material response. Specimens were subject to axial and shear frequency sweeps from 10 Hz to 0.01 Hz and at temperatures of 4, 25, and 40C. During the axial sweeps only an axial stress was applied and during the shear sweeps, beside a small seating load, only shear stress was applied. Stress levels were selected in this study to be as small as possible to allow for nondestructive testing and for the material to remain in the elastic range [Alavia and Monismith 1994]. The United Sates Air Force also conducted research using the hollow cylinder device. This investigation was driven by rutting occurring in the asphalt concrete runways of F-15 jets. A 39

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hollow cylinder specimen was used in the laboratory study of asphalt concrete subjected to realistic (but very slow moving) vehicle load simulation. A statistically significant difference is illustrated between the permanent strain resulting from laboratory testing with no principal-plane rotation and that resulting from testing with principal plane rotation. The tests conducted with the hollow cylinder were able to identify statistically significant differences between two levels of wheel loading, two aggregates, and two asphalts [Crockford 1993]. Another researcher today is using the hollow cylinder to obtain fundamental properties of asphalt mixtures such as creep compliance and tensile strength at low and intermediate temperatures. Hollow specimens used in this research have a six-inch (150mm) outer diameter and a 4-inch (100mm) inner diameter, giving a wall thickness of one inch (25 mm). Hollow specimens were created by drilling a hole through a gyratory compacted solid specimen. In this test, pressure is applied to the inner cavity resulting in tangential tensile stress or hoop stress developing in the wall. This test hopes to become a surrogate test for the Superpave Indirect Tensile Test (IDT) in asphalt mixture design [Buttlar et al. 1998]. The stress states that can occur in the hollow cylinder are almost unlimited when one has the ability to vary the internal and external chamber, the axial, and torsional stresses. The HCT device allows for the rotation of principal stress and planes not allowed in other laboratory tests, and can apply internal and external pressures unlike the SST. The following equations help to illustrate the variation in stresses across a hollow specimen when subject to stresses in three directions. These stresses are based on the theory of linear elastic thick-walled cylinders. All the notations in the equations are based on Figures 2-7 and 2-8, with stresses are assumed to be uniform along the entire length of the sample. 40

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rriro p o Pi Figure 2-7. Plan view of a hollow cylinder with outer radius, r o and inner radius, r i and subject to varying outer (P o ) and inner pressure (P i ). r 0 r imaxHMAS p ecimen l Torque at Rotation Ri g idl y Fixedat Bottommax(r)maxrl== Single Shearing W r Figure 2-8. Profile view of a hollow cylinder specimen. 41

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The following equations are for the radial ( r ) and tangential ( t ) at any radius (r) across the specimen subject to different internal (p i ) and external pressures (p o ) based on the outer (r o ) and inner (r i ) radii [Singer and Pytel 1980]: 222222222222222222rrrpprrrrprprrrrpprrrrprpriooioiioooiitiooioiioooiir The average radial and tangential stresses across the specimen can be calculated and reduced to the following equations: ioiiootioiioorrrrprpaveragerrrprpaverage The location of the average stress is based solely on geometry and occurs at: oiaveragerrr The shear stress () varies linearly across the specimen. The following equation is for the shear stress at any radius (r) in the specimen [Ni 1987]: 2)(44maxmaxiooorrJJrTrrr The average shear stress across the specimen is given as: 42

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223332ioioeqeqaveragerrrrrJTr The vertical stress is assumed to act uniformly over the specimen; the equation for the vertical stress is given below [Hight et al. 1983]: 222222ioiiooiovertivalrrrprprrW As shown above stress states across the specimen are not uniform. A hollow cylinder specimen should be thin enough to allow a uniform of stress distribution, but thick enough to contain a representative sample and provide for the maximum particle size. The length should be adequate to minimize end effects. Most guidelines established for hollow cylinder testing have been based on soil testing. For instance, as a rule of thumb, it has been suggested to have a wall thickness to grain size ratio of 10 to 25 when testing in the hollow cylinder [Hight et al. 1983]. This is to ensure a representative sample. This is not a tremendously difficult when testing clays or fine sands but unreasonable for testing asphalt concrete where aggregate sizes may approach one inch. Past researchers have used a one-inch wall thickness when testing asphalt concrete with maximum aggregate sizes of 3 / 8 to 3 / 4 (9.5 to 19 mm) [Buttlar et al. 1998 Alavi and Monismith 1994]. Guidelines have been established for the length and wall thickness based on equations from the theory of thin elastic cylinders [Saada and Townsend 1981]. (A thin walled cylinder is one where the stress distribution across the wall can be considered uniform). These equations were established to ensure that normal stress non-uniformity across the sample is kept to a minimum. However, these equations were based on the assumption of a thin cylinder. This 43

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assumption is reasonable for soil testing where the hollow cylinders could be considered thin. If a hollow cylinder is assumed to be thick, these equations break down and actually suggest that a very thick specimen is just as desirable as a very thin specimen. A finite element analysis is typically employed [Hight et al. 1983, Vaid et al. 1990, Buttlar et al. 1998] to quantify stress concentrations and non-uniformities for different geometries and provide insight into end effects. The finite element analysis provides boundary conditions not possible through the theoretical equations that assumed free boundaries and a thin cylinder. The issue of shear stress non-uniformity across the specimen does not depend on weather one considers a thin or thick specimen since the shear stress varies linearly across the specimen. The maximum shear stress occurs on the outside edge and the minimum shear stress occurs on the inside edge. The level of shear stress non-uniformity across a specimen is typically quantified with the following non-uniformity coefficient [Timoshinko and Woinowsky-Krieger 1959]: dr1rr1oirravgavgio3 One researcher has suggested a 3 value of less than 0.11 is acceptable [Vaid et al. 1990]. The important issue, as with any laboratory test, is to consider the actual stress state on the specimen by understanding the geometric configuration of the specimen and the boundary conditions. Summary The discussion presented indicates there are many laboratory tests today that are currently being used to evaluate a mixtures performance in the field. Some are large-scale and others small-scale. Some induce many stress states on the asphalt concrete laboratory specimen and others attempt to induce a single stress state on the asphalt laboratory specimen. The hollow 44

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cylinder device is a test that can achieve multiple stress states and stress paths as well as principal plane orientation not capable in other laboratory tests. 45

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CHAPTER 3 HEAVY VEHICLE SIMULATOR TESTING ON ASPHALT PAVEMENTS The opportunity to analyze the progression of instability rutting under controlled field conditions became a reality when the Florida Department of Transportation (FDOT) along with the South African Council of Scientific and Industrial Research (CSIR) purchased a Heavy Vehicle Simulator (HVS) Mark IV [Kim 2002]. Dynatest, a South African company, manufactures the Mark IV. Figure 3-1 The Mark IV Heavy Vehicle Simulator device. The Accelerate Pavement Testing (APT) facility was constructed in Gainesville, Florida, to test the performance of asphalt mixtures using the HVS. The insight gained by this testing and analysis was valuable in understanding the phenomenon of instability rutting. 46

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Instability Rutting in Modified and Unmodified Pavements Under HVS Loading A testing program to evaluate the performance of modified and unmodified asphalt mixes under HVS conditions was instituted by the FDOT. Modified asphalt mixes refer to asphalt mixes where the binder used contains a polymer modifier. The polymer-modifying agent employed in this study was styrene-butadiene-styrene (SBS). The HVS employs a super-single radial tire with an average contact stress of 115 psi and a footprint of 12 inches wide by 8 inches long. The load is applied uni-directionally at a speed of 6 mph. The testing was performed at a uniform pavement temperature of 50 C made possible by an environmental control chamber. The modified and unmodified asphalt mixtures used in the study were both fine graded SP-12.5 mixtures, with the unmodified binder rated 67-22 and the SBS-modified binder rated 76-22 (Superpave nomenclature). The pavement system at the APT facility consisted of a 4-inch layer of asphalt concrete, a 10.5-inch limerock base, a 12-inch stabilized limerock subgrade, and a natural sandy subgrade (A-3 soil). The asphalt concrete layer was placed in 2 two-inch lifts with initial air void contents of 7%. A detailed description of the testing program and results has been published [Tia et al. 2002, Tia et al. 2003, Byron et al. 2003]. Rutting Depths Rut depths were measured by a laser profiler. The laser profiler consists of 3,538 lasers attached to a movable frame. Each laser is accurate to within 0.1 mm. The profiler has 59 rows of lasers in the longitudinal direction (parallel to the direction of travel) spaced every 4 inches with each row consisting of 60 lasers spaced every inch in the transverse direction (perpendicular to direction of travel). Laser profile measurements were taken incrementally after a certain number of passes from the HVS. A transverse profile of movement of the pavement surface, i.e. rutting, was then constructed. A straight-line was drawn over the surface profile tangent to the 47

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highest points. The greatest distance from the straight-line to the trough was taken as the rut depth. This method is known as the surface profile method and is the method most used in defining rut depth [Kim 2003]. Measurements were generally taken after every 100 passes up to 1,000 total passes, then every 500 up to 3,000 total passes, then every 1,000 to 5,000 until 100,000 passes or the rut depth was deemed sufficient. The results of the modified versus unmodified mixes at 50C demonstrated that modified mixtures do a better job at resisting instability rutting than unmodified mixtures. Five sections were tested and the descriptions of the each section are described below. Table 3-1. Test sections in HVS testing. Section Test Temperature (C) Mixture Type Symbol in Figures 1B 50 Modified 2B 50 Modified 4A 50 Unmodified 4B 50 Unmodified 5A 50 Unmodified Sections 4A, 4B, and 5A (unmodified mixture sections) had about two and half times the rut rate as compared with that of sections 1B and 2B (modified mixture sections). 48

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02468101214161820050000100000150000200000250000300000Number of PassesRut Depth (mm)4A5A4B2B1B Figure 3-2. Rut depth progression for the various mixes The results indicate that both the modified and unmodified sections have an early response (fist 1,000 passes) that is very similar. After about 1,000 passes, the sections begin to display differences. The modified sections rut rate decreases and achieves a stable linear progression. The unmodified sections rut rate continues at the same rate and does not reach a stable rut rate seen in the modified sections. Trench cuts taken after testing (Figures 3-3 and 3-4) had been completed displayed no base or subgrade deformation, indicating true instability rutting. 49

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Figure 3-3. Trench cut on one of the sections tested during the HVS study. Figure 3-4. Transverse cross-section of the asphalt pavement after testing. 50

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Rut Propagation Analysis through Surface Profiling The laser profiler provided detailed measurements of the surface movement of the pavement section. Figure 3-5 displays an average transverse profile of one of the sections after 100 and 1,000 passes and Figure 3-6 displays the range of the longitudinal measurements that were averaged to get a single point in Figure 3-5. -4.0-3.0-2.0-1.00.01.02.0-30-20-100102030Average Transverse Cross-Section Profile (inches)Deformation (mm) 1,000 Passes 100 Passes Transverse Location Investigated Figure 3-5. Average transverse profile created by averaging the 58 longitudinal measurements from the laser profiler. 51

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-4.0-3.5-3.0-2.5-2.0-1.5-1.0-0.50.004080120160200240Longitudinal Cross-Section (inches)Deformation (mm) Average Value Reported for Transverse Profile Figure 3-6. The 59 longitudinal values averaged. The apparent variability in the longitudinal profile was not significant since the scale is in millimeters. The slight bulge in the middle of the wheel-path in Figure 3-5 is not surprising and is often seen in granular material under footings. The bulge occurs since the material in the center of the loaded area is under confined vertical compression and cannot deform laterally, thus the vertical movement is limited primarily to one-dimensional vertical movement. With the high degree of definition of the surface deformation from the laser profiler, a careful analysis of the progression of the rutted surface was initiated. From each average transverse profile, a ratio was calculated of the total surface area elevated over the total surface area depressed after each measurement. This ratio would help to determine the differences in instability rut progression and response between modified and unmodified mixtures. These measurements are only based on surface changes in height and depth and cannot reflect the 52

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changes in densities or air voids. Figure 3-7 displays the method used in calculating the surface Figure 3-7. Average transverse profile with ratio e area ratio from each transverse rut measurement. quation. Area 1), elevated (Areas 2 and 3), and uts. pled -4.0-3.0-2.0-1.00.01.02.0-30-20-100102030Average Transverse Cross-Section (inches)Deformation (mm) Average TransverseProfile after 1,000Passes Area 1 Area 2 Area 3 Ratio Used in Presentation:[(Area 2+ Area 3)/Area 1]*100% Figure 3-7 displays the deformation of the depressed ( naffected outside areas of the transverse cross-section from the laser profiler measuremenFigure 3-8 displays the ratio of the surface area elevated to surface area depressed for five sections presented above. The ratio was calculated for every rut measurement -cumulativerather than incrementally. (Although rut rate could have been analyzed as well in a similar manner, assurance that the laser profiler was in the exact location for each measurement couwith very small incremental changes made the degree of error greater than the degree in accuracy and thus was not attempted.) The logarithmic scale was selected only because it leads to a straight-line plot. 53

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01020304050607080901001000100001000001000000Pass #Area Ratio (%)4B4A5A1B2B 100 Figure 3-8. The ratio of area elevated to area depressed for all 50 C sections. The ratio of elevated surface area to depressed surface areas is roughly the same for both the modified and unmodified sections in the first 100 passes, but the unmodified sections diverge from the modified sections at 200 passes for Sections 4A and 4B and at 500 passes for Section 5A. The ratio of the area elevated to area depressed remains constant throughout testing for the modified mixtures at around 40%. The ratio of the surface area elevated to surface area depressed for the unmodified mixtures increases rapidly as the testing proceeds, reaching almost 90% for Sections 4A and 4B at 100,000 passes. These results may suggest that rutting in the modified sections was primarily due to vertical densification where the material within the wheel path is primarily compressed with some material laterally displaced. The rutting in the unmodified sections was due to vertical densification but also significant lateral shoving/displacement. This hypothesis that the instability rutting in the modified sections was primarily vertical densification and the unmodified sections was primarily lateral shoving/displacement was investigated further through the analysis of asphalt cores taken after the HVS testing. 54

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Volumetric Analysis of Asphalt Cores The wheel-p ath was 16 inches wide 12-inch tire width plus 4 inch tire wander. Cores ches from the travel centerline), the elevated region (14 toan al alt were extracted from the wheel-path (5 in 16 inches from the travel centerline), and from 24 inches outside the travel centerline -area deemed undisturbed by the rutting -from Sections 1B and 5A. Figure 3-9 displays the fintransverse rut profile of Sections 1B and 5A after 110,000 and 100,000 passes, respectively. lift) by the FDOT. The following table summarizes the air void content and heights of the cores. -10-8-6-4-20246-30-25-20-15-10-5051015202530Average Transverse Rut Profile (inches)Deformation (mm)5A1B 8 Figure 3-9. Final transverse rut profile of sections 1B and 5A. The each core was divided into Lift 1 (first or bottom asphalt lift) and Lift 2 (top asph 55

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Table 3-2. Summary of asphalt cores from pavement sections 5A and 1B post-testing. Height per Region (mm) Air Voids per Region (%) Section Layer Wheelpath Elevated Outside Wheelpath Elevated Outside Lift 2 41.3 46.2 45.3 4.46 8.39 6.75 Lift 1 43.9 51.5 49.7 2.64 4.88 4.25 5A Total/Average 85.1 97.7 95.0 3.52 6.54 5.44 Lift 2 37.9 40.7 39.6 5.14 7.02 6.67 Lift 1 48.2 51.9 51.1 3.77 4.70 6.07 1B Total/Average 86.1 92.6 90.7 4.37 5.72 6.33 The change in total height from the wheelpath to the outside and from the elevated region to the outside from Section 5A is negative .9 mm (downward movement) and +2.7 mm (upward movement), respectively. These values can be seen in Figure 3-8 and shows that the wheelpath core was likely taken 4 inches to the Right of the Centerline and the elevated core was likely taken 14 inches to the Right of the Centerline. The change in total height from the wheelpath to the outside and from the elevated region to the outside from Section 1B is negative .6 mm and +1.9 mm, respectively. These values can be seen in Figure 3-8 and shows that the wheelpath core appears to have been taken from approximately 4 inches Right of the Centerline and the elevated cores appear to have been taken from approximately 14 inches Right of the Centerline the approximate same offsets as the cores from Section 5A. The approach taken to clearly and concisely illustrate the primary mode of instability rutting each section underwent was through phase diagrams based on the following concepts. 56

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Each core is assumed to be composed of two volumetric phases a solid phase and void (air) phase. The solid phase consists of the aggregate and binder and the void volumetric phase consists of air. Air void content (AV%) can therefore be defined below as 100VVvAV% where V is the total volume, or the sum of the V v (air) and V solid and For each core, the volumetric phases can be simplified into heights the cross-section area is the constant of all the cores. Therefore, the total height (H) of each lift consists a certain height of air (Ha) and solids (Hs), or as follows: H = H a + H s And therefore, AV% = Ha/H. Ha Hs H Figure 3-10. Phase diagram of the asphalt cores. With this simple concept and the air voids and heights of the cores, an analysis of the cores can be undertaken clearly and simply. The Table 3-3 provides the heights and change in heights for each phase of each of the lifts of Sections 1B and 5A. 57

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Table 3-3. Summary of core data through use of phase diagrams. Section Lift Phase Initial Heights (mm) Wheelpath Region Heights (mm) Phase Change of Total Height Change Elevated Region Heights (mm) Phase Change of Total Height Change Air 3.06 1.84 30% 3.87 96% Two Solid 42.26 39.42 70% 42.29 4% Air 2.11 1.16 16% 2.52 22% 5A One Solid 47.61 42.72 84% 49.03 78% Air 2.64 1.95 40% 2.86 20% Two Solid 36.98 35.94 60% 37.83 80% Air 3.10 1.81 44% 2.44 -79% 1B One Solid 47.98 46.37 56% 49.48 179% The percentages reveal the story. If the entire change in height change in the wheelpath was due to vertical densification, then the percent loss attributed to air would be 100%; and vice-versa if the entire change in height was due to shoving then the percent loss attributed to solid loss would be 100%. (The percent loss of air and solid must add to 100%.) The wheelpath cores for Section 5A indicate that only 16% and 30% of the height change was due to air void loss (densification) with 70% to 84% due to solid material loss (shoving) in Lifts 1 and 2, respectively. However, the cores from Section 1B indicate an almost an even 50/50 split in the loss of height due densification and shoving. The percentages from the elevated region present a much more interesting phenomenon. Section 5A in Lift 2 indicates very little gain in solid material (4%), but a significant gain in air (96%). Section 5A Lift 1 the gain in height is primarily due to solid material (78%). Section 1B 58

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Lift 2 is very similar to Section 5A Lift 1 with 78% of the height gain coming from solid material. However, Lift 2 actually loses air voids with the infusion of solid material (-79%). This suggests that the elevated region in both sections gain in height by material being shoved away from the wheelpath. Once the material is displaced laterally from Section 5A (unmodified section) the material appears to expand or possibly dilate, while in Section 1B (modified section) the expansion or dilation of this material is not a significant contributor to elevation in height rather. One-way Versus Two-way Directional Loading The FDOT also conducted an evaluation on the most efficient method of HVS loading. This study investigated one-way versus two-way directional tire loading. The results clearly showed that one-way directional loading resulted in faster rut progression and in greater rut depths. The uni-directional loading caused the rut to develop at a rate of approximately 65 percent greater than that of the bi-directional loading. The results of this study are presented below. 59

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Figure 3-11. Comparison of one-way and two-way directional loading with no tire wander. Figure 3-12. Comparison of one-way and two-way directional loading with tire wander. It is important to note that the one-way loading sections had slightly lower average temperatures than the two-way loading sections [Tia et al. 2001, Tia et al. 2002, Byron et al. 2003]. Differences in rutting from one-way or two-way directional loading have not been 60

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observed at low temperatures [Huhtala, M. and J. Pihilajamaki 2000]. It is important to note that the rutting that occurred was indeed instability rutting. Trench cuts conducted after the one-way and two-way directional loading showed that there had been no deformation of the base layer; rutting was confined to the asphalt layer only i.e., true instability rutting similar to Figure 3-3 and 3-4. Figure 3-13 below illustrates directional axes inserted for illustrative purposes and to define terms used below. igure 3-13. Instability rutting with directional axes. ons. First, it is well known that asphalt concretup Z YX F The directional testing results raised two questi e exhibits visco-elastic properties [Ullidtz 1987]. In two-way directional loading, the relaxation time is about half that of the one-way directional loading (tire load must be picked and moved back to the starting position after each pass in one-way directional loading). A greater 61

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relaxation time should provide time to recover more strains, but actually the one-way directional loading with greater relaxation times between passes produces the greater rutting. (Not to mention that the one-way directional loading sections had an average temperature lower thatwo-way directional loading sections). The following figure illustrates the additional time that is required during one-way directional loading compared to two-way directional loading. n the econd, if instability rutting is thought to be due to shoving near the surface caused ges of the wheel-path (Huang 1993, Drakos VS. Two-Way Loading One-Way Loading Figure 3-14. Two-way versus one-way directional loading. S primarily by transverse shear stresses (Plane YZ) near the ed et al. 2001, and Huber 1999], then there should be little difference in the rut depths between one and two-way directional loading. Since the stress paths in the YZ-Plane will not be different from one-way and two-way directional loading. This is described in Table 3.4 (theletters A, B, and C, will be used for illustrative purposes). 62

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Table 3-4. Deformation of an element in an asphalt pavement near the edge of the tire in the YZ-plane (transverse plane). Shape of Element in AC Layer, Left Side of Tire Edge Point of Interest Shape of Element in AC Layer, Right Side of Tire Edge Far Away, Approaching (A) Next to, (B) Far Away, Passed (C) The above table illustrates that shear stresses that cause the deformation will not discern between one-way and two-way directional loading. Two passes of the wheel load in one-way loading will produce a shear deformation pattern of A-B-C A-B-C; two passes of the wheel load in two-way loading will produce a shear deformation pattern of A-B-C C-B-A. The magnitude of B would not vary between passes or upon the one-way or two-way directional loading. Stresses in the longitudinal plane (plane XZ) must also have an affect on the degree of rutting since these stress paths along this plane will vary depending on whether the load application one-way or two-way directional loading. The deformation of an element in the asphalt pavement near the tire edge in the XZ-Plane (longitudinal Plane) is illustrated in Figure 3-15. 63

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Figure 3-15. Deformation of an element in an asphalt pavement near the edge of the tire in the XZ-plane (longitudinal plane). For a cycle of two passes in two-way loading, an element on the edge of the wheel path in the asphalt layer will experience a pattern of shear stresses of: A-B-C-D-E E-D-C-B-A. For a cycle of two passes in one-way loading, an element in the asphalt layer will experience a pattern of shear stresses of: A-B-C-D-E A-B-C-D-E. The magnitudes of the shear stress are equal and opposite for positions B and D, as would be expected. It is not just shear stress reversal, from positive to negative, but the shear stress cycle reversal from positive to negative to negative to positive (two-way directional loading) versus positive to negative to positive to negative (one-way directional loading) that may be the cause for the difference in rut rate and depths between one-way and two-way loading. Quantification of Shear Stresses Within Asphalt Pavement To answer the above questions, the stresses that occur within the pavement in the area thought to be the most critical to the propagation of instability rutting top 2 inches near the tire edge [Dawley 1990 and Sousa et al. 1991] would need to be quantified. This was accomplished through three-dimensional finite element modeling of a radial tire contact stresses on the pavement system used during the FDOT study. The details of the methodology of the finite 64

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element analysis are presented in Chapter 4. The results are briefly summarized here specific to the HVS testing but were also consistent with the result sin Chapter 4. The HVS used a super-single radial tire with an average contact stress of 792 kPa (115 psi) and a footprint of 30.5 cm (12 inches) wide by 20.3 cm long (8 inches) (Byron et al. 2003). Since the exact tire contact stresses were not known for the radial tire used in the HVS, typical tire contact stress data on a nine-inch wide super-single radial tire rated at 792 kPa (115 psi) was available and was used in the analysis. Tire contact stresses were based on a tire contact measurement system developed by M.G. Pottinger [Pottinger 1991], which was especially developed for tire research, and consists of 1200 distinct measurement points, which register contact stresses in the x, y and z directions. This resulted in over 3,600 distinct stress measurements for the nine-inch wide super-single tire radial tire with five ribs and a gross contact area of 300 cm2 (47 in2) with an inflation pressure of 792 kPa (115 psi). The measurements provided a high definition of actual tire contact stresses. These stresses were then applied as the surface loading to the pavement system in the three-dimensional finite element analysis. The final three-dimensional mesh consisted of 204,185 nodeswith three degrees of freedom per node, resulting in a total of 612,555 degrees of freedom. The elements under the radial contact area had uniform dimensions of 7.62 mm by 10.16 mm (0.3 by 0.4 inches). Table 1 lists the elastic moduli, Poissons ratio and layer thickness of the pavement layers used in the analysis. 65

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Table 3-5. Properties of pavement layers used in the finite element analysis. LAYER Modulus (MPa) Poissons Ratio Thickness (cm) Asphalt Concrete 2,071 0.45 10 Base 193 0.45 26 LR Stabilized Sub-Base 128 0.45 30.5 Subgrade 55 0.45 116 The modulus for the asphalt concrete was determined from IDT resilient modulus tests at 25C on the same unmodified asphalt concrete used in the testing close to the average ambient testing temperature in the directional loading analysis [Tia et al 1999]. The base and subgrade values were based on LBR values conducted by the FDOT and converted to modulus values using the 1986 AASHTO formula [AASHTO 1986]. The Poissons ratio was selected to ensure minimal volumetric changes, to simulate a moving tire load during the shear-driven phase of instability rutting. As previously mentioned, the area in the asphalt layer believed most critical to instability rutting is the upper two inches near the edge of the tire. It was at the location depicted in Figure 3-14, that the shear stresses were culled form the Finite Element Analysis. 66

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AC Layer Base Subgrade Ti r e L o ad Critical Location Critical Location Y Z X Figure 3-16. Area of pavement system where shear stresses were obtained. Figures 3-17, 3-18, and 3-19 display the three directional shear stress magnitudes (YZ, XY, and XZ planes) in the longitudinal wheel travel directional from the finite element analysis along the edge of the tire at depths of 1.0, 2.5, 4.0 cms (0.5, 1, and 1.5 inches). Idealized depictions of the shear deformation for an element in the asphalt layer under the shear magnitudes reported are presented as well. 67

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-300.0-250.0-200.0-150.0-100.0-50.00.050.0100.0150.0200.0250.0300.0-30.0-20.0-10.00.010.020.030.0Longitudinal Position, Distance Off Centerline (cm)ShearStress (kPa) 1.0 2.5 4.0 Direction of TravelACB Figure 3-17. Shear stress in YZ-plane. -300.0-250.0-200.0-150.0-100.0-50.00.050.0100.0150.0200.0250.0300.0-30.0-20.0-10.00.010.020.030.0Longitudinal Position, Distance Off Centerline (cm)Stress (kPa) 1.0 2.5 4.0 Direction of TravelCABDE Figure 3-18. Shear stress in XZ-plane 68

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-300.0-250.0-200.0-150.0-100.0-50.00.050.0100.0150.0200.0250.0300.0-30.0-20.0-10.00.010.020.030.0Longitudinal Position, Distance Off Centerline (cm)Stress (kPa) 1.0 2.5 4.0 Direction of TravelACDEB Figure 3-19. Shear stress in XY-plane. The maximum shear stress in the longitudinal XZ-plane, seen in Figure 3-18, produces shear stresses about one-half of the maximum shear stress in the transverse YZ-plane, shown in Figure 3-17. The maximum shear stress in the XY-plane, shown in Figure 3-19, is about one-third of the maximum shear stress in the transverse YZ-plane, shown in Figure 3-17. It also interesting to note that the maximum shear stresses occur at depths of 4 cm and shear stresses do not vary much in the top 4 cm. To summarize, transverse shear stresses, shear stresses in the YZ-plane, near the edge of the wheel-path will undergo shearing only in one direction as the wheel approaches and passes -increasing in magnitude as the wheel approaches and decreasing in magnitude as the wheel passes as shown above in Table 3-4. This stress pattern of A-B-C C-B-A in two-way loading and A-B-C A-B-C in one-way loading does not explain the difference in rutting depths between one-way and two-way loading since the magnitude of B is the same. 69

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However, the stress pattern in the XZ-Plane and XY-Plane from one-way versus two-way directional loading will vary. One-way loading produces a stress pattern of A-B-C-D-E A-B-C-D-E while two-way loading produces a stress pattern of A-B-C-D-EE-D-C-B-A. Since the stress s pattern in these planes is what differs between one-way and two-way loading, the answer to why one-way and two-way loading produce different rut rates and depths may lie here. Analytical Approach To determine whether the lack of the shear stress cycle reversal, as seen in one-way loading, has an effect on permanent deformation, a simple analytical model was designed. Asphalt concrete is a complex material that seldom exhibits solely elastic deformation. Deformations in asphalt concrete typically contain viscous, visco-elastic and plastic deformations in addition to the elastic deformations. A simple model that can employ visco-elastic deformations, including permanent creep deformations, is the Burgers Model, also known as the four-element model [Ullidtz 1987]. The Burgers Model has the advantages of capturing instantaneous and delayed elastic deformation responses upon loading as well as instantaneous, delayed, and permanent deformation responses upon unloading. The Burgers Model is simply a Maxwell and a Kelvin Model connected in series, as shown in Figure 3-20 [Finedley et al. 1986]. Figure 3-20. Burgers model. 70

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The constitutive equation for the Burgers Model is described below. Four properties are required, two elastic constants representing the elastic modulus (R1 and R2) and two viscous constants (h1 and h2). The total strain at time t will be the sum of the strains from the spring and dashpot in the Maxwell Model and the strain in the Kelvin Model, as written below: 321 where 1 is the strain in the spring: 11R and 2 is the strain in the dashpot, s is the applied stress: 12 and 3 is the strain in the Kelvin part: 23213R These all can be combined to give the following constitutive equation: RRRRRR22112121222111 The simplest way to obtain a solution is to use Laplace Transforms and their inverse. This has the advantage of simplicity and consistency. Applying the Laplace Transform to the above equations reduces them to algebraic expressions and transforms them into a function of the complex variable s instead of time (t) indicated by a caret (^): sRssRRsRRR2221122121222111 71

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The inverse Laplace transform will bring the above expression back into terms of t. The Burgers Model was implemented into a MathCAD worksheet, representing asphalt concrete, to determine whether the reversal of the shear stress pattern (two-way loading) over a series of two cycles results in differences in total strain as compared to no shear stress pattern reversal (one-way loading). The finite element stress path evaluation showed that an element in the upper inch of the asphalt layer does not experience shear stresses until the tire is 20.3 cm (8 inches) away. The maximum stress achieved in the asphalt layer along the edges of the tire in the longitudinal plane is 103 kPa. Assuming that the 20.3 cm (8 inch) HVS radial tire used in the testing is traveling at the rated constant speed of 9.6 kph (6 mph), neglecting the deceleration and acceleration times for slow down and speed up when reversing direction, and the fact that the test track is thirty feet long, an element would experience a loading for about 0.22 seconds. The relaxation time between passes for the one-way loading would be 6.6 seconds and for the two-way loading would be 3.2 seconds. A sinusoidal load, similar to the actual semi-triangular loading pattern found in the finite element analysis, with a peak stress of +103 and kPa (+15 psi and -15 psi) was applied for 0.22 seconds. The next loading, representing the next pass, will either be +103 psi then kPa for the one-way directional loading or -103 and +103 kPa for the two-way directional loading, each with their respective relaxation times, as shown below. 72

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-125-100-75-50-250255075100125CycleStress (kPa) 6.4 Seconds120.22 Seconds0.22 Seconds Figure 3-21. Longitudinal shear stress pattern in one-way loading. -125-100-75-50-250255075100125CycleStress (kPa) 3.2 Seconds120.22 Seconds0.22 Seconds Figure 3-22. Longitudinal shear stress pattern in two-way loading. 73

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Properties for the Burgers Model were obtained from IDT creep tests (Roque and Butlar 1992) on a typical fine-grade mix found in Florida similar to the one used in the HVS directional testing and are presented in Table 3-6. Table 3-6. Values used in Burgers model to represent asphalt concrete pavement. Property Value R 1 2.5 Gpa 1 86.33 GPasec R 2 0.5 GPa 2 50 GPasec The results from the analysis of the loadings with the Burger Model are displayed in Table 3-7. Table 3-7. Results of Burger model under two-way and one-way loading. Loading Method Strain (mm/mm) Two-way (Shear Cycle Reversal) 5.434 X 10 -8 One-way (No Shear Cycle Reversal) 3.219 X 10 -6 The results indicate that the one-way directional loading produces more strain than the two-way directional loading in the one-dimensional Burger Model. The difference in strain for the one-way directional loading was two orders of magnitude greater than the two-way directional loading. Although the magnitude of the predicted deformations is different to that seen in the field, the differences match the observations that one-way directional loading produces more deformation than two-way directional loading. Laboratory tests on stress cycle reversal produce modulus values of one-half to two-thirds when there is no stress cycle reversal [Kallas 1970]. This suggests that shear stress reversal, even with greater relaxation times, produces greater deformation as seen in the HVS testing. 74

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Summary and Conclusions The FDOT testing program conducted focused on the effects of binder modification and one-way versus two-way loading. However the response of mixtures under-going instability rutting under these tests was insightful. Analysis of the transverse surface profile and the density-height changes of the cores before and after testing indicate that the modified sections rutted primarily due to densification within the wheel-path. The analysis suggests that rutting for the unmodified sections was due primarily to material being displaced or shoved from the wheel-path. The lateral shoving and possible expansion of the material once displaced increased the rut depth measurement and would affect ride performance. Instability rutting therefore is a combination both densification and lateral shoving, with lateral shoving resulting in higher rut rates and depths. The modified sections can be said performed better than the unmodified sections for their ability to resist lateral shoving. A mixtures ability to resist lateral shoving caused by near-surface shear stresses will determine its resistance instability rutting. The one-way and two-way loading FDOT study focused on determining which method would produce rutting in the shortest period of time. The tests results showed that one-way loading produces greater rut rate and rut depths compared to two-way loading under similar conditions. The computational Finite Element Analysis, a simple analytical test, and an understating of the mechanics of the shear stresses occurring within the pavement indicated that simply viewing rutting as caused by transverse shear stresses is not sufficient. Instability rutting is truly a three-dimensional phenomenon, a complex problem requiring an investigation that does not begin overly simplistic. A test that can capture complex stress states and patterns, such as the Hollow Cylinder test, would be invaluable and an excellent starting point to begin the understanding of the mechanisms behind instability rutting. 75

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CHAPTER 4 THE APPROXIMATION OF NEAR-SURFACE STRESS STATES IN ASPHALT CONCRETE THROUGH THREE-DIMENSIONAL FINITE ELEMENT ANALYSES Background Finite element analyses have not been extensively used to model three-dimensional tire loadings due to the complexity of modeling a radial tire in three dimensions and the rather small (40 60 in 2 ) and highly non-uniform contact area. Moreover, typical pavement structures consist of a thin layer of asphalt concrete overlying a base course, which rests on the semi-infinite subgrade. To accurately model the non-uniform loading and provide adequate boundary conditions requires a large number of elements and the associated large amounts of memory are not available either on past or current PCs. Because of these limitations, one approach to model three-dimensional tire contact stress was to approximate the complex loading conditions with uniform circular loads, use a layered elastic solution program that solves for uniform circular loads, and apply the superposition principle [Drakos et al. 2001]. This was done using the computer program BISAR [de Jong et al. 1973]. BISAR allows for circular uniform loads with one stress in the normal direction to the pavement and another in shear on the pavement at a specific angle. The concept was that a series of small uniform circular loads of varying vertical and lateral stresses could represent the actual tire stresses. This produced valuable results demonstrating that radial tires produce high near-surface shear stresses not found in bias-ply tires. However, these results were based on averaging numerous measured contact stresses to produce a single uniform circular load. Many of these uniform vertical loads would then define the radial tire. Furthermore, abrupt changes in tangential surface stresses on a continuum can lead to infinitely high horizontal stresses affecting the realism of near-surface stress states [Soon et al., 2003]. A finite element analysis would 76

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provide for more accurate load applications with loads being directly applied as nodal forces and the use of non-linear elements would provide insurance against superfluous stress concentrations. The University of Florida recently purchased a Silicon Graphics Interface (SGI) multi-processor computer, also known as a super-computer. Because of the extensive memory and faster computing time than the average PC, the ability to use finite elements in the modeling of three-dimensional tire-pavement system could be revisited. In order to better understand the actual stresses that occur in an asphalt pavement from radial tires the commercial finite element code ADINA [Bathe 2001] was used in modeling the three-dimensional effects of measured tire contact stresses in a typical pavement configuration. All pavement layers were assumed to be linear elastic, and dynamic effects were ignored in favor of promoting a basic understanding of static stress states before complicating the analysis with dynamic effects. Due to the complicated nature of the measured radial tire contact stresses, contact surfaces were used extensively to control the size of the problem. The insight gained into the near-surface stress states believed to be influencing instability rutting from this three-dimensional analysis would be invaluable. Three-dimensional Finite Element Model The three-dimensional finite element model was constructed and analyzed using the commercial finite element code ADINA. The finite element model consisted of 30,204 27-node elements of varying dimensions. First, an assessment of the grid size requirements required to model an attire contact stresses in asphalt pavements was performed. A uniform stress distribution showed that stresses in the area of interest, near the tire footprint, were not significantly affected by the extent of the boundaries. The initial assessment demonstrated that the three-dimensional model should be at least 72 inches deep and extend laterally at least 60 77

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inches in each direction from the center of the tire contact load to adequately represent the semi-infinite half space conditions associated with pavement problems. Because of the size of this model, the refinement needed near the tire contact area, and the desire to remain within the 3:1 element length to width ratio, the resulting memory requirements for the Silicon Graphics multiprocessor computer available for this analysis exceeded the 1,300 MB RAM memory available. To overcome the limitations associated with building a traditional mesh, contact surfaces were introduced, where a fine graded mesh representing the loaded surface was attached (glued) onto a coarse-graded mesh. This allowed for the introduction of coarse meshes at distances further away from the loaded area where the change in stress was more gradual, and far field stresses dominated the response. The use of contact surfaces was further justified based on the primary area of interest being the near surface area under and immediately surrounding the loaded tire, thus negating any possible negative numerical effects of far away contact surfaces. The final three-dimensional mesh consisted of 260,455 nodes with three degrees of freedom per node, resulting in a total of 781,365 degrees of freedom. Elements were all 27-node brick elements. Twenty-seven node brick elements were selected over 8-node brick elements in that 27-node brick elements possess a linear distribution of strain rather than constant strain, are less stiff, and have a better rate of convergence [Cook 1995]. The elements under the radial contact area had uniform dimensions of 0.30 by 0.40 inch. Contact surfaces were used for the transition from the asphalt layer to the base, from the base to the foundation, and from the fine mesh near the tire contact area to the peripheral areas. The lateral boundaries were fixed in the X and Y directions, but free in the vertical, Z, direction. The bottom of the mesh was fixed in all directions. The mesh itself extended 60 inches laterally in each direction and 72 inches vertically. 78

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Figure 4-1 illustrates the final three-dimensional mesh, with Figure 4-2 showing a plan view of the contact area of the three-dimensional mesh. Figure 4-1. Three-dimensional finite element mesh used in the analysis. Figure 4-2. Plan view of the three-dimensional finite element mesh used in the analysis. 79

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Pavement Structure and Loading Conditions Analyzed Pavement Structure A typical three-layer (asphalt concrete, base, and subgrade) pavement structure was used in this analysis [Myers et al. 1999, Drakos et al., 2001]. The asphalt layer had a thickness of 8 inches. The base layer was 12 inches and the subgrade was 52 inches. The properties of each layer were assumed to be isotropic, homogenous, and linear elastic. Table 4-1 lists the elastic moduli, Poissons ratio and layer thickness of the pavement layers used in the analysis. Table 4-1. Material properties of the various layers used in the analysis. LAYER Modulus (psi) Poissons Ratio Thickness (ins) Asphalt Concrete 100,000 0.45 8 Base 40,000 0.45 12 Subgrade 15,000 0.45 52 A low modulus for the asphalt concrete was selected to correspond to a warm summer day for a fairly new pavement the most critical time for the onset of instability rutting -while base and subgrade values were typical of measured values in the State of Florida. The Poissons ratio was selected to ensure minimal volumetric changes, as would be expected from a single moving tire load. Loading Conditions Analyzed The following loading conditions were analyzed: 1. Radial Tires with an inflation pressure of 115 psi a. 5-ribs with a gross contact area of 48 in 2 (R299 tire designation) b. 4-ribs with a gross contact area of 52 in 2 (M844 tire designation) 2. Uniform Vertical Load of 115 psi with a gross contact of 48 in 2 Gross contact area for the radial tires refers to the entire area between the outer edges, not just the area in contact with the pavement. Radial tire stress measurements were obtained from Smithers Scientific Services, Inc., in Ravenna, Ohio using a method described in Chapter 2. Stresses and displacements were recorded every 0.20 inch longitudinally (parallel to wheel path) 80

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and every 0.15 inch transversely (perpendicular to wheel path. This resulted in 1,200 stress locations or 3,600 distinct stress measurements for the radial tire with five ribs and a gross contact area of 48 in 2 and 1,300 stress locations or 3,900 distinct stress measurements for the radial tire with 4 ribs and a gross contact area 50 in 2 The measurements provided a high definition of actual tire contact stresses. Figures 4-3 and 4-4 depict the variation in vertical stresses for the two radial tires, and Figures 4-5 and 4-6 display the variation in transverse shear stress. Figure 4-3. Variation in vertical stress across the 5-rib radial tire used in the analysis. 81

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Figure 4-4. Variation in vertical stress across 4-rib radial tire used in the analysis. Figure 4-5. Variation in horizontal stress across 5-rib radial tire used in the analysis. 82

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Figure 4-6. Variation in horizontal stress across 4-rib radial tire used in the analysis. The figures display the truly non-uniform loading conditions and the variation in stress intensity not only across the tires but also across each individual rib. Load Application The measured tire contact measurements reported above were uniform stresses acting over areas of 0.03 in 2 in size. Three different stresses were provided, namely vertical normal stresses, transverse shear stresses, and longitudinal shear stresses. Ideally, each uniform stress should be applied to a single element. Unfortunately, the number of elements needed would have again exceeded the memory of the Silicon Graphics multiprocessor computer available for the analysis. Thus, the use of fewer elements was required under the contact area, which subsequently required the determination of the equivalent nodal forces to be applied to each node. The appropriate nodal forces for each element were determined by converting each uniform stress into an equivalent concentrated force. These forces were then applied to each element and distributed over the element nodes according to the rules described below [Hughes 1987]. 83

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If it is assumed that the concentrated forces {f} has components f x f y and f z then the element load vector, {r e }, acting on a surface of an element, is defined by: dSNreSTe where [N] T is the transpose of the shape function matrix, and {} is the surface traction vector. The contribution of {f } to {r e } can be determined by viewing the concentrated force as a large traction, {}, acting over a small area, dS. Subsequently, the concentrated force vector {f } can be denoted as: {f } = {}*dS The integral of [N] T {}dS thus becomes [N] T {f}, resulting in Equation (1), with n concentrated forces, becoming: 1TneiiirN f where [N] i is the value of [N] at the location of {f } i Elements under the loaded area were selected to be 0.30 inch by 0.40 inch to meet the memory requirements. Each element thus had 24 (eight x, y, and z forces) point forces acting on it. These point loads were then converted to the appropriate nodal forces through the method described above. Finally, Figure 4-7 shows a plan view of the radial tire contact stresses applied as nodal forces onto the pavement model surface at locations consistent with five tire ribs. Figure 4-8 shows a cross-sectional profile of the tire contact stresses. 84

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Figure 4-7. Contact area and the radial tire nodal forces used in the pavement response analysis. Figure 4-8. Cross-sectional view of applied tire nodal forces used in the pavement response analysis. Axisymmetric model A two-dimensional axisymmetric model (Figure 4-9) was generated in ADINA to provide a comparison between the stresses induced by a circular uniform vertical load and the more complicated radial tire loading effects. Because of the symmetric nature of the problem, only one half of the loaded area is modeled. The finite element model is 72 inches tall and 30 inches wide, with a uniform vertical surface load of 115 psi, distributed over a radius of 4 inches. The elements used consist of 8-noded isoparametric elements, with 72 vertical rows of elements, each containing 99 elements, for a total of 7128 elements. The layer thicknesses and elastic properties are the same as in the three-dimensional finite element model, discussed previously. 85

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Figure 4-9. A cross-sectional view of the axisymmetric finite element mesh used for comparison purposes. To evaluate the results from the axisymmetric finite element model, a comparison was performed between predicted shear stresses at the edge of the loaded area and those obtained from a semi-analytic layered-elastic theory solution, using the program BISAR [de Jong et al. 1973]. Figure 4-10 shows that the shear stress predictions obtained with ADINA [Bathe 2001] and BISAR [de Jong et al. 1973] are very similar for the exact same loading conditions, meaning that the axisymmetric model adequately captures the loading response due to the circular uniformly loaded vertical load. 86

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0.00.51.01.52.02.53.03.50.05.010.015.020.025.030.035.040.045.0Shear Stress (psi)Depth (inches) BISAR ADINA Figure 4-10. Comparison of shear stresses under the edge of a two-dimensional axisymmetric circular uniform vertical load predicted with BISAR and ADINA. Three-Dimensional Solution Process There are four different types of solution schemes available in ADINA, namely, 1) direct solver, 2) sparse solver for very large problems, 3) iterative solver, for nonlinear problems, and 4) multigrid solver, for parallel processing solutions. The direct solver requires a large amount of storage and is not recommended for large three-dimensional models [Bathe 2001]. The multigrid solver is intended for large three-dimensional problems with very large systems of equations, but when contact surfaces are used, ADINA does not allow for the use of the multigrid solver. The best solver for large memory limited problems is the sparse solver. Unfortunately, for the number of equations anticipated, the system requires a 64-bit version solution solver, with only the 32-bit version currently available for this research. Hence, since the iterative solver is also recommended for large problems, it was used by default. The equilibrium equations to be solved in a non-linear static analysis of a finite element model with contact surfaces in ADINA are: 0ttttRF 87

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where t+t R is the vector of the external nodal loads and t+t F is the force vector equivalent to the element stresses at time t + t [Bathe, 2001]. In non-linear analysis three iteration methods/schemes are available in ADINA, namely, 1) full Newton method, with or without line searches, 2) modified Newton method, with and without line searches, and 3) the Broyden-Flecther-Goldfarb-Shanno (BFGS) matrix update method. In this study, the full Newton method with lines searches was employed, based on its ability to converge and obtain accurate solutions [Bathe 2001]. The iterative solution process required 1295 Mbytes of RAM memory, with a resulting solution time of 9,000 seconds, using a single processor. Results of Three-Dimensional Analysis The vales to be reported from the FEM analysis should be viewed as more qualitative than quantitative. The reason for this approach is that the asphalt pavement was modeled as a linear elastic material and the FEM analysis was not as robust in terms of the number and sizing of elements as was originally hoped. However, the results should provide insight into the stress states that occur within the asphalt pavement from radial tires that will help lead to a better understanding of the mechanisms behind instability rutting. As an initial evaluation of the three-dimensional stress states that occur within the pavement in areas where instability rutting is observed was undertaken. Stress states in the longitudinal direction direction of tire travel -were analyzed to compare the three-dimensional differences between radial tire loadings and uniform vertical loading. The location selected for a stress-path comparative analysis was a point in the asphalt pavement one (1) inch away from the loaded area and one (1) inch deep into the pavement. The stress states at this point was also studied 10 inches ahead and behind in the longitudinal direction (x-direction) of the loaded areas. 88

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25 mm 2 5 mm2 5 mm X Y Outer Rib ZY Figure 4-11. Location within asphalt pavement from finite element analysis for preliminary stress state evaluation. The maximum shear stress, the confinement stress, and the orientation of the principal stresses were analyzed. The maximum shear stress, max is defined as the difference between the first and third (largest and smallest principal stresses in magnitude, respectively) divided by two: max = ( 1 3 )/2 Confinement, p, is defined as the summation of the three principal stresses divided by three: p = ( 1 + 2 + 3 )/3 Presentation of the orientation of the principal planes was done by calculating the Lode Angle. The Lode Angle, indicates the orientation of the line from a three dimensional stress state from the hydrostatic axis. The Load Angle varies from +30 to -30. The +30 state indicates the stress sates are the principal stresses, with 1 > 2 = 3 which can be viewed as a state of triaxial compression. The -30 state indicates the stress sates are the principal stresses, with 1 = 2 > 3 which can be viewed as a state of triaxial extension. A Lode Angle of 0 would indicate a state of hydrostatic stress, i.e. 1 = 2 = 3 The Lode Angle was selected as an efficient means to depict the rotation of the principal planes in a complex three-dimensional 89

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problem and to highlight the differences between radial tire loadings and uniform loadings conditions. The following four (4) figures depict the maximum shear, the confinement, the stress-path, and Lode Angle in the longitudinal direction. An image of the outer loaded area (length of the loaded area is 6 inches) is depicted on the y-axis of the graphs to help visualize the position of the stress magnitude compared to the tire position. Although these stress states are from a static analysis, a stress path of the loading from a moving tire load also can be inferred from the images. -10-8-6-4-202468100510152025303540MAXIMUM SHEAR STRESS (PSI)LATERAL POSITION FROM TIRE CENTER (INCHES) RADIAL TIRE UNIFORM LOAD Figure 4-12. Maximum shear stress in longitudinal direction along outer tire rib. 90

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-10-8-6-4-20246810051015202530CONFINEMENT (PSI)X POSITION (INCHES) RADIAL TIRE UNIFORM LOAD Figure 4-13. Confinement stress in longitudinal direction along outer tire rib. 0510152025303540051015202530Confinement (psi)Maximum Shear (psi) Radial Tire Uniform Load Figure 4-14 Confinement and maximum shear stress path. 91

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-10-8-6-4-20246810-30.0-25.0-20.0-15.0-10.0-5.00.05.010.015.020.025.030.0LODE ANGLE (Degrees)LATERAL POSITION FROM TIRE CENTER (INCHES) RADIAL TIRE UNIFORM LO AD Figure 4-15. Lode angle in longitudinal direction along outer tire rib The non-uniform stresses caused by the radial tires configuration (a 5-rib M844 radial tire) induce approximately 40% more shear stress within the pavement than the uniform loaded area. The largest maximum shear stress occurs directly adjacent to the middle of the loaded area. In addition, the magnitude of the confinement of the radial tire is approximately 10% less than that of the uniform loaded area. This is due to the fact that the radial tires vertical load is distributed unevenly more vertical load is concentrated in the middle ribs and less at the outer edges. The Lode Angle orientation suggests that the lateral stresses induced by radial tire configuration produce more of an extension state of stress than a vertical uniform load. Since a Lode Angle of -30 was not obtained, a simple triaxial extension test will not be able to capture the principal plane orientation that is induced by radial tires. 92

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As was somewhat expected, the radial tire produced much higher ratios of shear to confinement than the uniform load condition. This analysis was conducted on linear elastic material; if a material with failure/yield criteria with pressure dependency was instead used to model the asphalt pavement, radial tires may induce stress states into the pavement that will result in stresses that may yield the material and will result in a re-distribution of stresses that are not predicted by a simplified uniform vertical load. It is important to note that the largest ratio for the radial tire occurs at the corner of the loaded area rather than adjacent to the middle of the tire where the maximum shear was calculated. Finally, the three directional shear stresses ( xy yz xz ) were compared to evaluate the contribution of each to the maximum shear stress. The shear stresses at the locations analyzed above are presented below in the following figure. -10-8-6-4-20246810-20-1001020304SHEAR STRESS (PSI)LATERAL POSITION FROM TIRE CENTER (INCHES) 0 SHEAR XY SHEAR YZ SHEAR XZ Figure 4-16. Three directional shear stresses in the longitudinal directional along the outer tire rib 93

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The largest shear stress occurs within the transverse plane the YZ-plane, which is where the maximum shear also occurs. This suggests the most critical plane for investigating instability rutting is the transverse plane adjacent to the center of the loaded area. Transverse shear stresses have generally been thought to be the primary mechanisms behind instability rutting [Drakos et al. 2001, Dawley et al. 1990]. However, reversal pattern of the other directional shear stresses, the ratio of maximum shear to confinement, and the results of the one-way and two-way loading studies, indicate that instability rutting is a true three-dimensional phenomenon where shear stresses in the non-transverse plane play a role. To simplify the further investigation, the research focused on the transverse plane as the critical plane which would turn the complex phenomenon into a two-dimensional study. The three-dimensional aspects of instability rutting can not be ignored. Maximum shear stresses within the transverse plane are presented along the middle transverse cut of the 5-rib radial tire, 4-rib radial tire, and uniform vertical loading conditions, respectively with directional arrows showing the direction of the smallest angle formed between the maximum shear stress and the horizontal. Figure 4-17. Maximum shear stress magnitude (in kPa) and direction under the 5-rib radial tire loading condition. 94

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Figure 4-18. Maximum shear stress magnitude (in kPa) and direction under the 4-rib radial tire loading condition. Figure 4-19. Maximum shear stress magnitude (in kPa) and direction under the uniform loading condition. A further investigation of stress states in this area revealed that there existed two critical locations were the maximum shear stress was the highest within the pavement area adjacent to the tire loading. Outer Tire Rib B A 0.3 0.5 1.15 1.375 Figure 4-20. Critical stress locations for high shear stress 95

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The following figures present through Mohr Circles the stress state at these critical locations for the three loading conditions analyzed. -60-40-200204060-20020406080100120140 Horizontal PlaneStresses Vertical PlaneStresses M844 R299 Uniform 1 = 56 psi 3 = -15 psi 3 = 52 psi 3 = -12 psi Figure 4-21. Stress states at point A for the three loading conditions analyzed. 96

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-60-40-200204060-20020406080100120140 Horizontal PlaneStresses Vertical PlaneStresses Uniform R299 M844 1 = 65 psi 3 = -12 psi 3 = -8 psi1 = 60 psi Figure 4-22. Stress state at point B for the three loading conditions analyzed. The figure clearly indicates that the normal and shear stresses induced by radial tires are driving the stress state into planes of tension compared to a uniform load. The figures also display that the principle stresses are in a state of extension -with extension defined as the first principle stress having an angle of less than 45 with the horizontal plane. A compression state of stress is defined as the first principle stress having an angle equal to or greater than 45 with the horizontal plane [Lambe and Whitman 1969]. Significance of High Shear at Low Confinement Asphalt mixtures are known to be pressure dependent materials and have been modeled as Mohr-Coulomb materials [Krutz and Sebaaly 1993, McLeod 1950]. This is significant because HMA mixture strength parameters are often obtained at high these higher confinements. The c 97

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and parameters when idealizing asphalt mixtures to a Mohr-Coulomb envelope are typically based on strength tests in these high confinement regions and extrapolated, but not measured, to areas of low confinement. At elevated temperatures, asphalt mixtures begin to behave like granular materials with little cohesion as the binder becomes less viscous. In addition, the envelope in these low confinement regions may be curved, as indicated by the dashed line in Figure 4-23. Typical strength envelopes in hot mix asphalt and granular materials. Figure 4-23, which is sometimes the case for granular materials [Lambe and Whitman 1969]. ilure circle with os -25-20-15-10-50510152025-20-1001020304Normal StressShear Stress Granular Material Asphalt Concrete at Normal Operating Temperatures Asphalt Concrete at High Temperatures 0 Even unconfined vertical loading conditions will result in a Mohr-Coulomb fa ne principle stress at 0, meaning material response is determined from stress states and principal plane orientations that do not appear representative of the stress conditions in the asphalt occurring from radial tire loadings. In terms of rutting, the material response in areathought to be the critical to instability rutting are under high shear and low confinement. 98

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Summary and Conclusions The shear stresses of the radial tires differ from the uniform load in both magnitude (up to 20% greater) and distribution. The stress distribution also differs between the radial tires. All three loading conditions indicate the formation of shear planes under the loaded area. The shear planes will tend to shove the material up and away from the tire resulting in rutting within the asphalt pavement. The shear stresses are more prevalent under the radial tires near the surface compared to the uniform loading. The three-dimensional analysis of measured radial tire contact stresses indicated that stress states in the asphalt layer are characterized by low levels of confinement, and even tension, coupled with high shear stress. In contrast to this, stress states induced by uniform vertical loading conditions, which are traditionally assumed in pavement analysis, are characterized by higher levels of confinement and lower shear. Furthermore, these more critical stress states from radial tire contact stresses occur in the near surface region, near the edge of the tire, where instability rutting is known to occur. Uniform vertical loadings do not produce these critical stress states in this region. The results presented also imply that the characterization of instability rutting requires testing at these low confinement (and sometimes tensile) stress states and principal lane orientation, rather than at the higher stress states typically used in the strength characterization of mixtures. Accurate modeling of radial tire contact stresses is necessary to describe the mechanism of instability rutting. It appears that an accurate description of mixture properties (shear strength and deformation responses) at stress states (high shear at low confinement) in the critical instability-rutting region (edge of the rib at shallow depths) may be necessary to properly evaluate a mixtures instability rutting potential. The next chapter investigates what laboratory 99

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tests would be available to produce similar stresses and orientations that were indicated in the FEM analysis as occurring in the asphalt pavement in these critical regions. 100

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CHAPTER 5 CREATION OF A HOLLOW CYLINDER DEVICE Introduction The computational analyses presented in Chapter 4 found radial tiers induce higher shear stresses at lower confinements than bias-ply tiers in areas of the asphalt pavement where instability rutting is observed to occur. It is believed that instability rutting is driven by high shear stresses at low confinements. The orientation of the principal planes induced by tire loads in these areas is not aligned with the vertical and horizontal planes, but orientated 20-30 from the horizontal. If these stress conditions are critical to instability rutting then it is of interest to test under these critical stress conditions. Testing under these conditions is believed will help determine a mixtures susceptibility to instability rutting more accurately than current tests which do not test under these conditions. The laboratory equipment currently available triaxial device cannot obtain these critical stress states orientation of the principal planes. Currently, two laboratory testing devices can match the stress states mentioned the true triaxial device and the hollow cylinder device. The true triaxial device is a complex, rare, research-only tool that has no feasible industrial or mainstream research capability. The hollow cylinder device, on the other hand, can be instituted with typical research testing equipment. Although, hollow cylinder testing is not typically used outside of research institutions it is more widely understood, recognized, and used at research institutes far more than the true triaxial device. For these reasons, a hollow cylinder testing device (HCTD) was constructed at the University of Florida to test asphalt mixtures under these critical stress states. This chapter describes the development of the HCTD at the University of Florida. 101

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Hollow Cylinder Specimen Dimensions The first task prior to developing the mechanical apparatus was to determine the dimensions of the hollow asphalt concrete (AC) specimen. Once the specimen dimensions were determined, the fabrication of the HCTD could proceed and an analysis of stress concentrations under typical testing conditions for the selected geometry could be undertaken. The dimensions of the hollow cylinder will affect the stress distribution across the specimen under axial and torsional loading. The wall thickness must be thick enough to accommodate the range of aggregate sizes in the asphalt mixture to ensure a representative sample but, be thin enough to allow for more uniform stress distribution of across the wall. In addition, means and methods to easily and readily construct the hollow cylinder must also be considered. A previous study investigating rutting induced by fighter jets on asphalt concrete runways employed a one-inch thick wall thickness [Crockford]. This study employed cyclic torsional testing with inner and outer pressure differentials. Present researchers also use a one-inch thick wall thickness in their HCTD for the investigation of low-temperature crack propagation in asphalt pavements [Buttlar] these studies do not employ torsion on the samples but a greater inner pressure than outer pressure to induce a tangential hoop stress. In Florida, SP-9.5mm and SP-12.5mm are typically the most common Superave asphalt structural mixes employed. These size mixtures would therefore be the most likely employed in pavement design and studied in pavement analyses. The largest aggregate size that would therefore be encountered with some frequency for these mixture designations is the nominal maximum aggregate size of 9.5 or 12.5 mm (a SP-9.5mm mixture has a maximum aggregate size of 12.5 mm and a SP-12.5mm has a maximum aggregate size of 19 mm.) The HCTD wall 102

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dimensions must be sized to ensure that a representative sample of a SP-9.5mm to a SP-12.5mm mix is present. The University of Florida employs a ServoPac Gyratory compaction device for creating laboratory asphalt concrete specimens. The ServoPac can produce either 100mm or 150mm diameter specimens. These diameter sizes are common for gyratory compactors used in asphalt research and industry. Specimens made to a 100mm diameter can be up to 150mm in height. Specimens made to a 150mm in diameter can be up to 135mm in height. Most guidelines established for hollow cylinder wall thicknesses have been based on soil testing. For instance, as a rule of thumb, it is recommended that the wall thickness should be 10 to 25 times the average grain size to ensure a representative [Saada and Townsend]. If one used this recommendation for asphalt concrete testing, the typical SP-12.5mm mixture has an average grain size (defined as D 50 ) of approximately 4mm, thus a wall thickness ranging from 40 mm to 100 mm would be recommended. (The largest particle size in a SP-12.5mm is 19mm.) Based on the above information, the outer diameter of the hollow cylinder specimens will either be 100mm or 150mm and the height either 150mm or 125mm for the respective heights. The inner diameter is the one parameter that must be designed for that best meets the desires of a representative sample and stress uniformity. An initial investigation looked at inner diameters in 25mm increments that would produce wall thicknesses of 25mm or greater. Table 5-1 indicates the possible specimen dimensions for the hollow cylinder. 103

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Table 5-1. Possible hollow cylinder specimen dimensions. Specimen Outer Diameter, d o (mm) Parameter 100 150 Height, L (mm) 150 135 Inner Diameter, d i (mm) 25 or 50 50, 75, or 100 Wall Thickness (mm) 37.5 or 25 50, 37.5, or 25 d i /d o 0.25 or 0.50 0.33, 0.50 or 0.67 L/d o 1.5 0.9 As mentioned in Chapter 2, there are guidelines, recommendations, and techniques that are available that can be employed to evaluate a hollow cylinder specimen. These are in place to help ensure that the hollow cylinder has the least amount of stress non-linearity and can be considered as representative sample. An analysis based on closed-form solutions was carried out to determine the effects that the different proposed sample geometries have on the distribution of the tangential, radial and shear stress across the wall of the proposed hollow cylinder geometries. Based on this analysis, the selection of geometry could be made based on knowledge of stress distributions, limitations, and deficiencies between the different geometries. A finite element analysis would then be employed to investigate end effects and stress concentrations on the proposed hollow cylinder geometry not possible from the closed-form solutions. A determination based on the quality of the results from the FE analysis would determine whether the geometry could be implemented for design and production. Closed-form analysis of stresses across hollow cylinder wall The level of stress non-uniformity of the radial and tangential stresses across a hollow cylinder specimen due to different internal and external pressures was undertaken. Each of the seven possible specimen geometries had an inner and outer pressure applied to achieve an 104

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average uniform radial stress, r of 20 psi and an average uniform tangential stress, t of 30 psi. The average stress and the inner and outer pressures stresses were calculated from equations found in Chapter 2. Table 5-2. Inner and outer pressures applied to proposed hollow cylinder geometries. d i (mm) d o (mm) d i /d o Inner Wall Pressure, p i (psi) Outer Wall Pressure, p o (psi) 25 100 0.25 5.0 23.8 50 100 0.50 15.0 22.5 50 150 0.33 10.0 23.3 75 150 0.50 15.0 22.5 100 150 0.67 17.5 21.7 The quantitative values of the inner and outer diameters do not play a role in this analysis, only the ratio of inner diameter to outer diameter (di/do). Figures 5-2 and 5-3 display the non-linearity of the radial and tangential stress distributions across the wall of the hollow cylinder specimen for the different specimen geometries (note that r is defined as the radius from the center minus the inner radius, r i ). The stresses have been normalized, so a value of one represents the average stress, either 20 psi for the radial stress or 30 psi for the tangential stress. 105

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00.20.40.60.811.21.400.10.20.30.40.50.60.70.80.91r/(ro-ri)r(r)/r,average di/do = 0.25 di/do = 0.33 di/do = 0.50 di/do = 0.67 Figure 5-1. Ratio of radial stress to average radial stress across the wall of the specimen for various geometries. 00.20.40.60.811.21.41.600.10.20.30.40.50.60.70.80.91r/(ro-ri)t(r)/t,average di/do = 0.25 di/do = 0.33 di/do = 0.50 di/do = 0.67 Figure 5-2. Ratio of tangential stress to average tangential stress across the wall of the specimen for various geometries. When a torsional stress is applied to a hollow cylinder specimen, the shear stress varies linearly across the wall thickness. The maximum shear is on the outside edge, while the 106

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minimum shear is on the inside edge. Each specimen was evaluated by determining the slope of the shear across the wall when a torsional stress is applied. Figure 5-4 displays these results. Again, the shear stress has been normalized based on the average shear stress occurring on the specimen based on Equations from Chapter 2. 0.000.200.400.600.801.001.201.401.600.00.10.20.30.40.50.60.70.80.91.0r/(ro-ri)(r)/average di/do = 0.25 di/do = 0.33 di/do = 0.50 di/do = 0.67 Figure 5-3. Ratio of shear stress to average shear stress across the wall of the specimen for various geometries. Table 5-3 gives the stress non-uniformity coefficient 3 from Chapter 2 for shear stress across the specimen for the different possible geometries. The values reported parenthetically are the percentage above the recommended value of 0.11 [Vaid et al. 1983]. Table 5-3. Beta 3 shear stress non-uniformity quantities for the various geometries. d i /d o 3 0.25 0.28 (155%) 0.33 0.24 (118%) 0.50 0.16 (45%) 0.67 0.09 The results from Figures 5-2 to 5-4 indicate that a thinner wall produces a greater uniformity of stress across the hollow specimen than a thick wall, as would be expected. This 107

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difference is markedly seen when comparing either a di/do of 0.25 or 0.33 to 0.67, but not as great when comparing a di/do of 0.50 to 0.67 for the radial and tangential stress distribution. Thus, a di/do of 0.67 is not seen as a great improvement on a di/do of 0.50 for radial and tangential stress distribution. The 3 parameter is significantly affected by wall thickness. At this point a decision of the geometry of the hollow cylinder specimen became clearer, but still engineering judgment would have to be employed. The specimens with outer diameters of 100mm were excluded for the following reasons: 1. An ID of 25 mm resulted in too much stress non-uniformity 2. An ID of 50mm (25mm wall thickness) would produce a average grain size ratio of an SP-12.5 to wall thickness below 10 3. Latex membranes to seal the inner radius of a 100mm hollow cylinder are not readily available (The outside of the hollow cylinder as well as the inside of the specimen (pi and po) could potentially be under different pressures. Water or air could be used as the pressuring medium; if water is used then asphalt specimen must be protected against water infiltration.) Specimens with an outer diameter of 150mm would have to have an inner diameter approximately between 75mm to 100mm. The improvement in stress uniformity for radial and tangential stress is not significant, but for shear it is. However, a 25mm wall thickness was considered more of a concern lack of a representative sample and grain size to wall thickness ratio -than a high 3. Based on the above information, desire for a representative sample, and engineering judgment, a specimen size with an outer diameter (OD) of 150 mm, an inner diameter (ID) of 70mm, and a height of 135mm was selected. The specimen size would provide a wall thickness of 40mm meeting the recommendations for hollowing cylinder testing of soils by Saada for an average SP-12.5mm mixture and helping ensure a more representative sample. Figure 5-5 displays the specimen dimensions selected that will be used for the HCTD: 108

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ID = 70mm OD = 150mm Wall Thickness = 40mm Height = 135 mm Figure 5-4. Hollow cylinder specimen dimensions. A finite element analysis was now employed to investigate end effects and stress concentrations on the proposed hollow cylinder geometry not possible from the closed-form solutions. A determination based on the quality of the results from the FE analysis would determine whether the geometry could be implemented for design and production. 109

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Finite element analysis of stress non-linearity and concentrations The purpose of the finite element analysis was to detect and evaluate stress concentrations that occur for the selected geometry as well for any coring eccentricities that may occur. The finite element program ADINA [Bathe 2001] was employed to model the different specimen geometries under different loading conditions. The analysis employed a linear elastic material model with a modulus of 100,000 psi and a Poissons ratio of 0.35. The mesh consisted of 832 twenty-seven-node brick elements. The bottom boundary was fixed in all directions. The top boundary was free to move only in the z-direction for the radial/tangential and vertical stress analysis. For the torsional stress analysis, the top of the specimen was kept in plane and a displacement was prescribed resulting in 0.01% strain. The analysis was three-dimensional. Figure 5-5. Three-dimensional finite element model of the hollow cylinder specimen for end effects analysis. Again, an outer pressure of 22.5 psi and an inner pressure of 15 psi was placed on the specimen. This, according to the closed-form solution, would produce an average stress of 20 psi in the radial direction and 30 psi in tangential direction. The closed form solution predicts a 110

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maximum radial stress equal to the outer and inner pressures at the associated boundaries, i.e., a maximum radial stress of 22.5 psi on the outer wall and 15 psi on the inner wall. The closed form solution predicts a maximum tangential stress of 35 psi on the inside wall and a minimum tangential stress of 27.5 psi on the outside wall. Figure 5-6. Tangential (hoop) stress distribution across hollow cylinder wall from finite element analysis. 111

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Figure 5-7. Radial stress distribution across hollow cylinder wall from finite element analysis. Both figures display stress concentrations near the ends of the hollow cylinder specimen from the loading condition. It appears that the middle third is relatively little affected by the end restraints for the radial stresses, but the tangential stress actually is uniform. The tangential stress distribution along the middle third, although affected by the end constraints in that it doesnt match the closed-form solution, may actually be considered an improvement since its now appears uniform. For a closer inspection of the analysis, the stress distribution form the closed-form analysis was compared across the wall at three locations from the finite element analysis. Figure 5-8 depicts the heights along the hollow cylinder where the stress distribution across the wall was compared to the closed-form solution. 112

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135 mm 22.5 mm 45 mm 67.5 mm Figure 5-8. Three vertical locations where stress distributions across the wall from the finite element solution were compared to the closed-form solution. 10.0015.0020.0025.0030.0035.0040.0000.10.20.30.40.50.60.70.80.91r/(ro-ri)tangential Closed Form ADINA -67.5mm ADINA -45 mm ADINA -22.5 mm Figure 5-9. Comparison of the tangential stress across the hollow cylinder wall from the closed-form solution and the finite element analysis with end constraints. 113

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10.0012.0014.0016.0018.0020.0022.0024.0000.10.20.30.40.50.60.70.80.91r/(ro-ri)radial Closed Form ADINA -67.5mm ADINA -45 mm ADINA -22.5 mm Figure 5-10. Comparison of the radial stress across the hollow cylinder wall from the closed-form solution and the finite element analysis with end constraints. From Figure 5-9, the middle 45mm appears to have a relatively uniform tangential stress distribution. The uniform tangential stress is near to the average tangential stress of 30 psi calculated form the closed-form solution. From Figure 5-10, the middle third and some more closely matches the closed-form solution. The end constraints caused stress concentrations only in top and bottom third as shown by Figures 5-9 and 5-10. Thus, an ID of 70 mm on a 150mm OD sample can be considered acceptable in terms of stress distribution. The three-dimensional results in Chapter 4 indicate that the normal longitudinal stress (the second principal stress) is approximately 60% the transverse normal stress (the first principal stress). As previously stated in Chapter 4, the transverse plane within the pavement next adjacent to the loaded area is assumed to be the critical plane for instability rutting. Instability rutting is primarily the effect of stresses within the transverse plane both shear and normal stresses acting within the transverse plane reducing the problem effectively to two-dimensions. In addition, testing equipment limitations did not allow for testing under different inner and outer 114

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pressures at the time of this work, requiring the inner and outer pressures remaining the same during testing -thus eliminating a hoop stress being induced in the specimen and reducing the problem to twp-dimensions. Based on the previous analyses, applying no differential pressure between the inner and outer pressures will only reduce stress concentration and stress discontinuities across the specimen creating a more uniform stress within the specimen. Future testing that employs differential inner and outer pressures will likely require additional investigation into stress discontinuity due to loading conditions. Creation of a Hollow Asphalt Concrete Specimen Hollow asphalt concrete specimens can be created through methods that require unique molds, methods, and machining [Sousa 1988]. The selection of the University of Florida hollow cylinder was based a desire to ensure that specimens could be easily and quickly produced with existing production methods and machinery. The method selected, that is both effective and efficient, is to core out the center of a laboratory compacted solid cylindrical specimen (or a field core). Asphalt concrete cylinders are easily and routinely created using gyratory compactors. A coring drill is also available. Coring bits typically come in quarter-inch diameter increments producing a diameter (inner diameter of hollow specimen) of an additional four-hundredth of an inch. Currently, another researcher employs this method for the production of hollow cylinders [Buttlar 1998]. Figure 5-11 depicts the hollow cylinder specimen production concept. 115

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Solid Cylindrical Asphalt Concrete Core Out Center with Drill Hollow Cylindrical Asphalt Concrete Specimen and Solid Core Figure 5-11. Hollow cylinder specimen production. It has been well documented that gyratory compacted samples do not have uniform density. Specimens compacted in the gyratory compactor have higher density (lower air voids) in the center and lower density (higher air voids) near the top and bottom and along the cylinder edges. This is more pronounced in taller specimens and wider specimens [Harvey et al. 1994, Masad et al. 1999, Tashman et al. 2001]. Other researchers who employ coring for hollow cylinder production have not found density non-uniformity to be such an issue [Buttlar 1998]. Nonetheless, air voids of the solid cylindrical specimen prior to coring and the post-cored solid center can be calculated using ASTM D1189 and D2726. From these quantities, a reasonable estimate of the air void in the hollow specimen can be calculated through volumetric averaging. Hollow Cylinder Specimen Production Aggregate Preparation and Batching The first step in production is to obtain and procure the aggregate as described below: Virgin aggregate obtained from the field is dried in the oven at a temperature of 300F for twelve hours and then allowed to cool at room temperature until it can be handled. The virgin aggregate is then sieved and separated into individual particle sizes based on Superpave sieve sizes (19 mm), (12.5mm), 3/8 (9.5mm), #4, #8, #16, #30, #50, #100, #200, and pan (minus #200). 116

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The aggregate material is then batched according to the Job Mix Formula (JMF) into 5000g samples. Mixing The next step in specimen production is to combine the aggregate with the asphalt binder is a process known as mixing described below: The batched aggregate (5000g), asphalt binder, and mixing equipment (pans, buckets, mixing tools) are placed in an oven for two hours at 300F. The appropriate amount of asphalt binder based on the Superpave design is added to the batched aggregate and thoroughly mixed until the aggregate has been fully coated with the binder. Short Term Oven Aging (STOA) and Compacting Once mixing is complete, the material is aged and compacted as described below: The mixed asphalt concrete is spread evenly in a shallow pan and placed in an oven at 275F for two hours. After one hour, the mixture is stirred to ensure uniform aging. This process is known as STOA. After the STOA, the mixture is then placed in the 150mm diameter Servopac gyratory mold (which as been heated at 275F for two hours) and compacted at 1.25 gyratory angle with 600kPa ram pressure for the appropriate number of gyrations to achieve 7% air voids. The compacted cylindrical specimens are extruded from the mold and allow to cool for 24 hours. The cylindrical specimens are then sawed on each side. The height of the cylindrical sample after sawing is 135mm. The Bulk Specific Gravity (G MB ) of the sawed specimens was then taken in accordance to ASTM D1189 and D2726. The percent air voids was then calculated from the G MB and the G MM of the specific mixture. STOA is used for instability rutting research since instability rutting is a phenomenon that often affects pavements in their first-year of service. 117

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Coring The cylindrical specimens were then cored to produce the hollow cylinder specimens as described below: The sawed cylindrical specimens were then cored down the middle using a rotary drill (Figure 5-12) Figure 5-12. Coring process to obtain hollow cylinder specimen. The Bulk Specific Gravity of the remaining core, is then calculated using ASTM D1189 and D2726. The percent air void of the hollow can then be calculated from the G MM With the air void content of the original cylindrical specimen and the core known, a reasonable estimate of the air void content of the ring can be calculated using volumetric averaging. The preceding procedures were repeated whenever a hollow cylinder specimen was produced. 118

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Hollow Cylinder Testing Device A GCTS hydraulic system and load frame is one mechanical apparatus currently employed at the University of Florida. The GCTS hydraulic system has the ability to apply both axial force and torque to material specimens soil, rock, concrete, or asphalt. The maximum axial force is kips and 500 in-lbs torque. The torque is applied through a horizontal hydraulic actuator and is controlled either a load cell (torque mode) or LVDT (displacement mode). A piston-cylinder constructed at the University of Florida, provides confinement to the sample chamber and is controlled by the use of calibrated pressure transducers. The piston-cylinder device has been integrated into the Testar IIm software. (The piston-cylinder uses water as the fluid medium.) Thus, axial forces, torsional forces, and confinements can be applied to material specimens simultaneously under pre-defined failure or time limits. The GCTS system is controlled through the MTS Testar IIm software. Data acquisition is conducted through the Testar IIm software as well. The user has the ability to input any test procedure with the appropriate data acquisition for a given time, number of cycles, or a determined response is reached. The GCTS hydraulic system has been used extensively at the University of Florida for research in asphalt materials and a more detailed description of the GCTS system and load frame can be found there [Frank 1999, Swan 1999, and Pham 2000]. Prior asphalt concrete testing at the University of Florida consisted of solid 100mm diameter specimens 150mm in height. Previous asphalt concrete material testing at the University of Florida with the GCTS used 100 mm diameter and 150mm high solid specimens. These specimens were tested in a chamber connected to the GCTS load frame and controlled by the Testar IIm software via a personal computer. Minor modifications were necessary to incorporate the GCTC system for hollow cylinder testing. The same Plexiglas chamber that had 119

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been previously used for 100mm specimen testing was employed for the hollow specimen testing. The components of HCTD, proceeding from bottom to top of the physical device, are outlined below with a figure below each section: 1. PEDESTAL. A hollow stainless steel Pedestal provides the foundation for the HCTD. The Pedestal is bolted to the chamber base through two (2) pairs of anchor bolts on flanges at the base of the Pedestal. The Pedestal has an inner diameter of 35mm. One (1) 5mm diameter opening was threaded through the Pedestal to allow the pressure on the outside and inside of the specimen remains the same. (This 5mm hole is threaded and sized for a screw-plug, such that the screw plug can be inserted if a pressure differential up to 20 psi is desired.) The Pedestal contains four (4) equally spaced threaded bolt-holes at the top to secure the bottom End Cap. Figure 5-13. Plan view of the pedestal. 120

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Figure 5-14. Side view of pedestal. 2. END CAPS a. Bottom End Cap. A hollow stainless steel Bottom End Cap provides the connection from the asphalt specimen to the Pedestal. The Bottom End Cap is two-tiered. The lower tier has a 6.9-inch outer diameter and the upper tier has an outer diameter of 59 inches. The lower tier forms a flange for the bolts to connect to the Pedestal. There is an O-ring groove on the End Cap side in contact with the Pedestal. The 5.9-inch diameter side of the Bottom End Cap is in contact with the AC specimen. This side is serrated at 20 lines/inch to provide additional frictional with the asphalt specimen. The asphalt specimen is glued to the serrated end through a high strength Epoxy. 121

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Figure 5-15. Plan view of the bottom end cap. Figure 5-16. Side view of the bottom end cap. b. Top End Cap. A hollow stainless steel Top End Cap provides the connection from the asphalt specimen to the Connection Cap (described below). The Top End Cap has a 5.9-inch outer diameter w. The side of the End Cap in contact with the asphalt specimen is serrated at 20 lines/inch to provide additional frictional with the asphalt specimen. The asphalt specimen is glued to the serrated end through a high strength Epoxy. The side in contact with the Connection Cap, non-serrated side, contains an O122

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ring groove. The non-serrated side also contains four (4) equally spaced threaded bolt holes. One (1) 5mm diameter opening was threaded through the Top End Cap to allow the pressure on the outside and inside of the specimen remains the same. (This 5mm hole is threaded and sized for a screw-plug, such that the screw plug can be inserted if a pressure differential is desired as with the Pedestal.) Figure 5-17. Plan view of the top end cap Figure 5-18. Side view of the top end cap. 123

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3. Connection Cap. The stainless steel Connection Cap serves as the link between the Top End Cap, and the Loading Arm of the GCTS system the load arm provides axial force and torque to the asphalt specimen. There are four (4) equally spaced smooth-bored bolt guide holes on the bottom of the Connection Cap that align with the four (4) threaded bolt holes on the Top End Cap. The top of the Connection Cap also contains a 13mm diameter groove, 10mm in depth acting as a guide hole for the GCTS Loading Arm. There are four (4) threaded bolt holes surrounding the groove. These threaded bolt holes connect the Connection Cap to the Stabilizer Collar. (The Stabilizer Collar was designed previously by others as a stabilizer and locking device for the GCTS Load Arm. The Collar was not modified for the HCTD but was utilized.) Figure 5-19. Plan view of the connection cap. Figure 5-20. Side view of the connection cap. 124

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Water is used as the medium for pressure control and, with an attached temperature control device, can be heated or cooled to allow testing under various temperature ranges. Latex rubber membranes (on the outside and inside the hollow cut-out) shield the asphalt specimen from the water. The inner membrane is held in place by the O-rings fitted between the Bottom End Cap and the Pedestal and the Top End Cap and the Connection Cap. A seal is ensured by the force exerted by the bolted connections. Epoxy is used to bond the asphalt specimen to the serrated End Caps through a method that has been successfully demonstrated previous solid torsional testing of asphalt concrete [Pham, 2003.] Top End Cap Connection Cap Bottom End Cap Threaded Opening Pedestal Figure 5-21. The hollow cylinder testing device. 125

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Stabilizer Collar Figure 5-22. Hollow cylinder testing device inside plexiglass chamber and surrounded by latex membrane. Summary The analyses presented in Chapter 4 found that significant shear stresses at low confinements are produced by radial tires in areas of the asphalt pavement where instability rutting is observed to occur. The orientation of the principal planes induced by tire loads in these areas is not aligned with the vertical and horizontal planes, but orientated 20-30 from the horizontal. The HCTD was created to tests laboratory prepared asphalt concrete specimens under these stress conditions. The next chapter details a testing program begun with the HCTD. 126

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CHAPTER 6 MATERIAL SELECTION AND TESTING PROGRAM Introduction The research approach taken thus far has been directed at the creation of a laboratory testing device that can more accurately replicate the stress states in the asphalt pavement that are believed to result in instability rutting. These concepts lead to the creation of the hollow cylinder testing device (HCTD). Material Selection To assess the ability of the HCTD to determine an asphalt mixtures susceptibility to instability rutting, two (2) asphalt mixtures with known susceptibility to instability rutting were selected for testing. Testing would be conducted on short-termed oven aged samples and at elevated temperatures (40C) to mimic the field conditions under which instability rutting is known to occur and propagate (new pavements during first hot season). One mixture was selected that performed well displayed little to no instability rutting and one mixture was selected that performed poorly displayed signs of instability rutting. Both mixtures were based on current Superpave mix design standards. The following table outlines the mixes selected for the testing program and general features of each mixture: 127

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Table 6-1. Superpave mixtures employed for testing program. Rutting Performance/Estimation Mixture Name Field APA IDT Superpave Mix Designation Aggregate Type Cu & Cc Unified Soil Classification System (USCS) of Aggregate White Rock Coarse 1 (WRC1) N/A Good Good SP-12.5mm (Coarse) Limestone 15.8 & 3.0 SW (3.6% Passing #200) Superpave Mix #1 (SP1) Poor Poor N/A SP-9.5mm (Fine) Granite 23.9 & 2.4 SW-SM (5.8% passing #200) The asphalt content of the WRC1 mix is 6.5% and the asphalt content of the SP1 mix is 5.5% -limestone is more absorptive than granite. PG 67-22 was used for all mixtures in this study. The Job Mix Formula (JMF) and grain-size distribution curves of the mixes can be found in the Appendix. The USCS designation was placed in the above table based on the assumption that at elevated temperatures, as the binder becomes less viscous; the aggregate frictional characteristics of the asphalt mixture play a larger role in shear resistance. The USCS was one method of presenting the aggregate characteristics of the aggregate portion of the mix. WRC1 is defined aswell graded sand; SP1 is defined aswell-graded sand with silt based on the USGS. The USCS designation should be viewed in conjunction with the Superpave mix designations in terms of aggregate size and distribution. Resistance to shear is hypothesized to be critical to a mixtures ability to resist instability rutting. (Both aggregates would be classified as A-3 sand based on the AASHTO soil classification system which is the standard system for roadway engineering.) 128

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Testing Program The purpose of the testing program for the HCTD was two-fold. The primary objective has been stated to qualitatively and quantitatively assess an asphalt mixtures susceptibility to instability rutting. The selection two mixes with different performances would provide a self-check of the results. The second objective was to ensure the ability of the HCTD to provide results that could be considered reliable. The ability of the HCTD to provide reliable results would investigate the affects of coring and spatial aspects of a hollow cylinder on stress-strain. The following tests and their objectives with the mixtures stated above were instituted: Table 6-2. Tests performed with the hollow cylinder test device. Test Objective Specimen Types Used in Testing Axial and Torsional Complex Modulus (G*) Affects of coring and spatial aspects of hollow cylinder specimens versus solid specimens at micro-strain levels Hollow Cylinder 1 100mm X 150mm Solid Cylinder Core 2 Extension Affects of coring and spatial aspects of hollow cylinder specimens versus solid specimens at small-strain levels Hollow Cylinder 1 100mm X 150mm Solid Cylinder Core 2 Cyclic Torsional-Extension Qualitatively and Quantitatively assess a mixtures susceptibility to instability rutting Hollow Cylinder 1 1 150mm Outer diameter, 70mm Inner diameter and 135 mm in height 2 ~68 mm diameter and 135 mm in height solid specimen the core remaining from the creation of the hollow cylinder 129

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The Chapter 7 describes the Axial and Torsional Complex Modulus (E* and G*) testing and the extension testing in more detail and presents the results of these tests; Chapter 8 details the Cyclic Torsional-Extension testing program and results. 130

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CHAPTER 7 COMPLEX MODULUS AND EXTENSION TESTING Complex Modulus Testing Background The complex or dynamic modulus test is a test that is being investigated as a test to predict the rutting susceptibility of hot mix asphalt mixtures by associating the asphalt mixtures elastic and viscous response at small strains to actual field rutting potential/susceptibility. Complex modulus testing (axial complex modulus testing) was first described as a test on hot mix asphalt in 1962 [Pupation 1962]. Recently, NCHRP Project 9-19 evaluated the complex modulus test and the AASHTO 2002 Design Project focused complex modulus testing as a Simple Performance Test (SPT) for the rutting resistance of HMA mixtures [Pellian and Witczak 2002]. The dynamic modulus test is outlined in ASTM D3497. In axial complex modulus testing, sinusoidal stress or strain amplitudes are applied axially to an unconfined cylindrical specimen at 16, 4, and 1 HZ. The ASTM Standard also recommends testing at temperatures of 5, 25, and 40C. The complex modulus test is similar to the unconfined creep testing in testing set-up, but the load application and analysis of the response differs. The dynamic or complex modulus test relates the cyclic strain to the cyclic stress in a sinusoidal load test. The complex modulus is defined as: 00* E where: 0 is the stress amplitude, 0 is the strain amplitude. 131

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The complex modulus, E*, is composed of a real component known as the storage modulus, E, and an imaginary component known as the loss modulus, E. The storage modulus represents the elastic portion of the response and loss modulus represents the viscous portion of the response. The storage and the loss modulus can be obtained by measuring the lag in the response between the applied stress and the measured strain. This lag in the response is known as the phase angle (). Torsional complex modulus testing is similar to axial complex modulus testing (ASTM D3497) but instead of applying a sinusoidal stress or strain amplitude in the axial direction, a torsional stress or strain is applied. The torsional complex modulus test at the University of Florida applies a cyclic torsional force to the top of specimen and measures the displacement on the outside diameter as shown in Figure 7-1. Knowing the torsional stress and strain, the shear modulus is then calculated based on the theory of elasticity. r 0r imaxHMASpecimen l Torque at peakRotation Rigidly Fixedat Bottommax(r)maxrl== Single AmplitudeShearing Strain Figure 7-1. Torsional shear test for hollow cylinder column. 132

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The dynamic shear modulus is calculated from the following relationship: *G where is the applied torsional shear stress and is the measures shear strain response. Assuming that pure torque, T, is applied to the top of a HMA column, the shearing stress varies linearly across the radius of the specimen. The average torsional shear stress, on a cross section of a specimen avg is defined as: avg = S/A where: A is the net area of the cross section of the specimen, i.e A = (r o 2 -r i 2 ), r o and r i are the outside and inside radius of a hollow specimen, respectively. (For a solid specimen, r i = 0), and S is the total magnitude of shearing stress. S can be calculated as: oirrrdrrS)2( where: r is the shear stress at the distance r from the axis of the specimen, i.e r = m r/r o where m is the maximum shearing stress at r = 0. On the other hand, the torque, T, can be calculated from: JrrdrrToirrmr)2( where: J is the area polar of inertia, J = (r o 4 r i 4 )/2. From above, m can be expressed as: m = Tr o /J 133

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From the above equations, one can write the equation for avg as: JTrrrrioioavg223332 or JTreqavg where: r eq is defined as the equivalent radius. It can be seen in above equation that r eq = 2/3r o for a solid specimen. r eq = 2/3 (r o 3 r i 3 )/(r o 2 r i 2 ) for hollow specimen. In practice, r eq is defined as the average of the inside and outside radii. Shear strain is calculated: lreq where: l is the length of specimen, and is the angle of twist. The angle of twist, can be measured either using an LVDT or a proximitor, The torsional complex modulus, G*, is composed of a real component known as the storage modulus, G, and an imaginary component known as the loss modulus, G. The storage modulus represents the elastic portion of the response and loss modulus represents the viscous portion of the response. The storage and the loss modulus can be obtained by measuring the lag in the response between the applied stress and the measured strain. This lag in the response is known as the phase angle (). In addition, Poissons ratio can be estimated when conducting axial complex modus testing and/or axial and torsional complex modulus testing. When conducting an axial complex modulus test lateral strains can be also be measured. With lateral and axial strains measured, the 134

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Poissons ratio can be determined. The shear complex modulus, G*, can be determined by the following equation [Harvey et al. 2001]: 21*EG This equation assumes that Poissons ratio is constant, although Poissons ratio has been seen as being frequency dependent [Sousa and Monismith 1987] and that linear elastic relations with moduli hold for visco-elastic complex moduli. Recent findings using the torsional complex modulus testing suggest that reasonable values of Poissons ratio can be determined using equation 7.10 from axial and torsional complex modulus testing -E* and G* [Pham 2003]. Testing Program Axial and torsional complex modulus tests were performed on the two (2) asphalt mixtures described in Chapter 6 SP1 and WRC1. Axial and torsional complex modulus tests were performed on both hollow and solid specimens. The hollow specimens were created in the manner described in Chapter 5. The solid specimens consisted of 100mm by 150mm specimens created in the ServoPac gyratory compactor and from the cores from the creation of the hollow cylinder. The axial and torsional complex modulus (E* and G*) tests were conducted using the GCTS load frame and at a temperature of 40C. The temperature control device uses water as the fluid medium for heating, thus all testing was performed under water with the specimens protected from the water by latex membranes. The three specimen geometries were prepared with the Superpave Gyratory compactor to air void content of 7.5%. The hollow cylinders air void content was calculated from volumetric averaging of the original 150mm diameter solid specimen and the leftover core specimen. The torsional displacement at the top of the sample was measured by fixing a two (2) small winged plates perpendicular to the Connection Cap refer to Chapter 5 with the plates 180 135

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from each other. Two (2) LVDTs were mounted to a fixed body -the support columns of the chamber and would measure the displacement of the winged plates and thus the specimen. The Testar IIm software collects the displacements from the two (2) LVDTs once testing has begun at a sampling rate set by the user. With the displacement measured, the known dimensions of the specimen and the torque being applied, the stress and strain can be calculated providing G*. For axial complex modulus testing, the winged plates were removed, and the LVDTs directly measured the axial displacement of the Top End Cap. Data reduction was performed by methods previously generated [Swan 2000 and Pham 2003]. Depending on the sample geometry (hollow, solid core, or the 100mm solid), the axial force and torque magnitude would vary to ensure a linear range was obtained for each specimen geometry. Due to experimental difficulties torsional complex modulus testing was not conducted on the solid core specimens for the WC1 and SP1 mixes. E* and G* Testing Results An analysis of the results was based on the assumption that E* and G* is a material property unique to each mixture at a given density. Thus, the means and methods in which a specimen is made or its geometry should not influence its E* or G* value for a similar density and any differences noticed between specimens may be assumed to be from specimen creation techniques. Results were analyzed qualitatively with emphasis of the response between the different methods and mixes and with little emphasis on the quantitative values. The following table presents the air void content for each of the three specimen geometries for the two mixtures. As previously mentioned, the air void content of the hollow cylinder was calculated by volumetric averaging of the original 150mm diameter sample (pre-coring) and the ~68mm diameter core (post-coring). 136

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Table 7-1. Summary of air void contents of different specimen geometries. AIR VOID CONTENT Specimen Geometry SP1 WRC1 Solid 100mm Diameter 7.344% 7.359% Hollow 7.294% 7.262% Solid Core (~68mm Diameter) 6.641% 6.592% 010020030040050060070080090010000200400600800100012001400160018002000Microstrain ()E* (MPa) WRC1 Hollow WRC1 Solid Figure 7-2. E* versus microstrain for different WRC1 specimen geometries under axial loading at 1HZ. 137

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02040608010012014016018020001002003004005006007008009001000Microstrain ()G* (MPa) WRC1 Hollow WRC1 Solid WRC1 Core Figure 7-3. G* versus microstrain for different WRC1 specimen geometries under torsional loading at 1HZ 0100200300400500600700800900100005001000150020002500Microstrain ()E* (MPa) SP1 Hollow SP1 Solid Figure 7-4. E* versus microstrain for different SP1 specimen geometries under axial loading at 1HZ. 138

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02040608010012014016018020001002003004005006007008009001000Microstrain ()G* (MPa) SP1 Hollow SP1 Solid SP1 Core Figure 7-5. G* versus microstrain for different SP1 specimen geometries under torsional loading at 1HZ. In general, the hollow cylinder and solid specimens had similar E* and G* values at higher microstrain levels and once the E*.G* to microstrain curves became more linear. The cores, as should be expected due to their denser state, had higher G* values that the hollow and solid specimens. The differences in complex modulus in specimen creation is much more pronounced at lower microstrain levels. More testing should be conducted to determine the influence of specimen creation on complex modulus response. However, since testing in this research for rutting potential will be conducted at higher strain levels, the observed similar response between the different specimen creation methods allows a level of confidence that testing at higher strain level does not warrant concern by the means and methods used to create the hollow cylinder specimen. A comparison between the hollow cylinder, solid, and core specimens from each mixture to each other is presented below in the following figures. 139

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0100200300400500600700800900100005001000150020002500Microstrain ()E* (MPa) WRC1 Hollow WRC1 Solid SP1 Hollow SP1 Solid Figure 7-6. WRC1 and SP1 hollow cylinder specimen E* results versus microstrain. 02040608010012014016018020001002003004005006007008009001000Microstrain ()G* (MPa) WRC1 Hollow WRC1 Solid WRC1 Core SP1 Hollow SP1 Solid SP1 Core Figure 7-7. WRC1 and SP1 specimen G* results versus microstrain. Evaluations As mentioned before these tests were designed for a qualitative evaluation and this was carried over when comparing the two different mixtures to each other. As indicated above, the SP1 mixture typically is stiffer at lower microstrain levels but decreases more rapidly such that at 140

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higher microstrain levels the WRC1 has higher G* and E* levels. The WRC1 mixtures also tend to have less of a dramatic decrease in stiffness as the microstrain level increases compared to the SP1 mixture. This suggests the SP1 mixture may be more of a brittle mix compared to the WRC1 mixture. Further complex modulus testing is recommend to determine the subtle difference between these two mixtures at these low strain levels. Extension Testing Extension tests were also performed on a WRC1 and SP1 solid, hollow, and core specimens as was done for the complex modulus testing. The purpose of extension testing was also primarily designed to qualitatively evaluate the different specimen creation techniques and compare the response of the two different mixtures. Extension testing was performed with the use of the GCTS system at 40C with both mixtures compacted to air void contents of 7.5%. The specimens were tested by increasing the lateral pressure on them such that there would be axial displacement. The pressure was increases at a rate of 1 psi every 30 seconds. The following p-q diagram displays the pressure that was induced on the specimens for the extension testing: 141

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02468101202468101214p (psi)q (psi) 16 Figure 7-8. P-Q diagram of stress path used for extension testing. The following two figures display the results of the extension testing of the different specimen configurations for the different mixtures: 00.511.522.50123456q (psi)Axial Strain WRC1 Core WRC1 Solid WRC1 Hollow Figure 7-9. Axial strain versus deviator stress, q, for different WRC1 specimen geometries. 142

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00.511.522.5012345Q (psi)Axial Strain 6 SP1 Core SP1 Hollow Figure 7-10. Axial strain versus deviator stress, q, for different SP1 specimen geometries. Due to equipment difficulties, extension testing was not performed of the SP1 100mm diameter solid specimen. Based on the results of the extension tests, the WRC1 specimens showed a very similar response. The SP1 hollow and core specimens differed to a greater degree than among the WRC1 specimens. In general, as expected due to lower air void, the core was the stiffest of the specimen geometries. In general, the responses indicate that different specimen geometries made through different means do not significantly affect the material response. Comparing the responses of the two mixtures against one another, the WRC1 appears to be more resistant to axial deformation than the SP1 mixture resulting in less axial strain for a similar degree of stress increase. This can be considered consistent with the E* testing, which indicated the WRC1 mixture was stiffer at higher microstrain levels. 143

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Summary and Conclusions Both mixtures specimen geometries displayed similar responses suggesting material properties are not dependent on the means and methods used to create testing specimens. In addition, the WRC1 mixture was seen to be a stiffer mixture than the SP1 mixture at higher strain levels. The next chapter seeks to analyze and evaluate the WRC1 and SP1 mixtures under conditions consistent with stress states found in the Chapter 4 analyses using the Hollow Cylinder Testing Device. 144

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CHAPTER 8 TORSIONAL EXTENSION TESTING Introduction This chapter discusses the response of cyclic torsional-extension tests on two Superpave asphalt mixtures. As previously mentioned, the two Superpave mixes were selected based on their ability to resist instability rutting WRC1 being the observed good performer and SP1 being the observed poor performer. The Finite Element analysis of Chapter 4 indicated that high shear at low confinement occurs in areas of the asphalt pavement where instability rutting is observed to occur when subject to radial tire contact loading. Thus, it was decided to conduct laboratory testing under high shear stresses at low confinement under cyclic loading conditions to simulate moving loads at 40C to simulate the stress states where instability rutting is observed to occur and under the most adverse environmental conditions for instability rutting. The response of each mixture under the above conditions would determine the ability of a laboratory test to determine a mixtures susceptibility to instability rutting. Further analysis of the response would possibly offer insight into the mechanisms behind instability rutting and/or possible future testing programs. The tests conducted also investigated the healing potential of mixtures under the above-mentioned conditions. Healing is a phenomenon seen in IDT resilient testing. Asphalt mixtures may heal or partially heal if loading is stopped for a period of time prior to the onset of macro-damage as if the asphalt mixture will respond as if it had never been loaded. Healing has been incorporated into the UF cracking model for asphalt pavements. The purpose of investigating healing under tests designed to investigate instability rutting was to identify the potential of an all-encompassing damage model that could incorporate both cracking and instability. 145

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Laboratory Loading Conditions The following figure displays, through the use of a Mohr-Circle, the critical stress state imposed on the samples during the cyclic torsional testing. -5-4-3-2-1012345-4-20246810NORMAL STRESS (psi)SHEAR STRESS (psi) Horizontal PlaneStresses Vertical PlaneStresses 1 = 5.7 psi3 = -1.2 psi POLE Figure 8-1. Stress state imposed on hollow cylinder samples during testing. The above stress states displays a shear stress of 3 psi, a lateral confining stress of 4 psi, and an axial confining stress of 0.5 psi on the hollow cylinder specimen. As can be seen, the stress state applied to the specimens is significantly less than the stresses calculated by the Finite Element analysis approximately 10 times less. The stress state values predicted by the finite element analysis ripped the asphalt specimens apart immediately at 40C after one load application cycle. A preliminary study found that the above stress state allowed for a linear response of the asphalt specimen until failure had been reached and was qualitatively similar to the stress state 146

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determined from the finite element analysis similar ratio of shear stress to lateral and axial stress. The fact that the quantitative values from the finite element (FE) analysis resulted in immediate failure reaffirms the conclusion of the author that the computational analysis was more of a qualitative assessment of the stress states in the asphalt pavement, rather than actual, and demonstrates the possible limitations of the FE modeling. As mentioned previously, tests were performed cyclically -haversine loading at 0.5 HZ with the stress state in Figure 8-1 representing the maximum stress state. Loading was increased from a hydrostatic stress state of 0.5 psi to the stress state displayed in Figure 8-1 in and then decreased to the hydrostatic stress state in a linear manner such that the Mohr-Circle grew at a uniform rate. No rest period was employed between each loading cycle. The following figure, in p-q space, represents the cyclic loading conditions imposed on the hollow cylinder asphalt specimens. -4-3-2-1000.511.522.5P (psi)Q (psi) Figure 8-2. Cyclic stress path imposed on asphalt specimens for testing. 147

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Laboratory Testing Program A series of two tests were designed. The first series of tests was to observe the overall response of the two mixtures under the torsional cyclic loading conditions and measure their overall response to each other. Second, each mixture would be tested to determine if the healing phenomenon occurs under cyclic loading conditions. Also, the magnitude of recovered strain was investigated during the rest period to determine the portion of elastic and plastic strain. The following table describes the tests performed on each of the mixtures under torsional cyclic loading: Table 8-1. Testing matrix for cyclic torsional testing. TEST CONDITION WRC1 SPECIMENS TESTED SP1 SPECIMENS TESTED Load to failure 2 2 Load to 90% failure, rest for 8 hours, re-load 1 1 Load to 50% failure, rest for 8 hours, re-load 1 1 Failure was defined as the state when rapid strain occurs resulting in the specimen being torn apart typically 5 to 7% radial strain and 3.5% axial strain. After radial/axial strains of this magnitude were reached, the specimen sheared and data collection was ceased. Laboratory Testing Results Testing to Failure The first series of tests were conducted on two (2) hollow cylinder samples of the WRC1 and SP1 mixes. Both samples were short-termed aged, had air voids of 7.0.5%, and were 115mm in height. The specimens were tested at 40C. Axial and shear movements of the top of the specimen were measured at intervals of 0.25 seconds. The axial and shear strains were then determined. In addition, the angle of axial to radial strain (), as an indirect measurement of the dilation/vertical expansion of the specimen, 148

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was also calculated; the larger the angle the greater the portion of axial strain to shear strain. This angle was considered insightful based on the response of the pavements in the HVS testing program of Chapter 3. The following figures display the radial and axial strain response of the two mixes and the calculated of the mixtures two specimens of each mix were tested, thus the fours lines on the graphs. 0123456789100200400600800100012001400160018002000CYCLESSHEAR STRAIN(%) (%) SP1_A SP1_B WRC1_A WRC1_B Figure 8-3 Shear strain versus number of cycles of two SP1 and WRC1 specimens under axial-torsional-extension testing. 149

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00.511.522.533.544.550200400600800100012001400160018002000CYCLESAXIAL STRAIN (%) SP1_A SP1_B WRC1_A WRC1_B Figure 8-4. Axial strain versus number of cycles of two SP1 and WRC1 specimens under axial-torsional-extension testing. 0510152025303540455055600200400600800100012001400160018002000CYCLESAXIAL/RADIAl ANGLE, (o) SP1_A SP1_B WRC1_A WRC1_B Figure 8-5. Angle of axial to shear strain, of two SP1 and WRC1 specimens under axial-torsional-extension testing. 150

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Both mixtures display a similar overall response in terms. This overall response can be divided into three ranges: Primary, Secondary, and Tertiary. The first range, or Primary Range, is the initial response and consists of large amounts of strain over approximately 50 to 100 load cycles due to the initial loading. After the Primary Range, the strain rate response (slope of curve) decreases until it reaches a minimum rate, around 200 load cycles. It now has entered the Secondary Range. The Secondary Range is marked by a stable strain rate. Finally, the strain rate begins to increase entering the Tertiary Range. The Tertiary Range is marked by increasing strain rate until failure/rupture is achieved. The following table gives the relative load cycles that the three ranges for the two mixtures based on axial, shear, and strain response. Table 8-2. Three ranges of the axial and shear strain response of the mixtures. Range WRC1 (Loading Cycles) SP1 (Loading Cycles) Primary 0 to ~200 0 to ~200 Secondary ~200 to ~1200 ~200 to ~1000 Tertiary ~1200 to 1400 (1) ~1000 to 1100 (1) (1) Cycles to Rupture When observing the radial strain response, both mixtures appear to be very similar, progressing at approximately the same radial strain rate. It is only when the SP1 mix breaks off earlier into its tertiary response does one observe the difference between the mixtures. The SP1 moves into its radial tertiary response at approximately 5% radial strain (950 to 1,025 cycles) while the WRC1 mixtures move into their tertiary range at approximately 6.5% shear strain (1,200 cycles). 151

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However, in the axial strain response, the difference between the mixtures is more apparent the SP1 mixture has a higher axial strain rate than the WRC1 mixture. It is when both mixtures reach approximately 2.5% axial strain do they move into their respective tertiary responses 900 to 1,000 cycles for the SP1 mixture and 1,200 cycles for the WRC1 mixture. The angle of axial to shear strain, versus number of cycles in Figure 8.5 demonstrates that the SP1 is undergoing more axial strain compared to shear strain than the WRC1 mixture. Load Relaxation at Tertiary Range, then Load Re-Application The next series of tests sought to investigate the phenomenon of healing. Loading conditions would be the same as before. The first set of tests focused on stopping the test when the SP1 and WRC1 specimens entered the Tertiary range --947 and 1,294 load cycles, respectively -then re-loading the specimens after a rest period of 8 hours. The stress condition during the 8-hour rest period would be an isotropic state of stress of 0.5 psi. The following figures present the results of these series of tests displaying axial and shear strains, and The re-loading was considered to be a continuation of the test beginning at the cycle which the initial test stopped. The strain presented for the reload tests is therefore a cumulative strain. 152

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0123456789100200400600800100012001400160018002000CYCLESSHEAR STRAIN (%) SP1_C WRC1_C SP1_Reload WRC1_Reload Beginning of Re-Load A pplication After Rest Period Figure 8-6. Shear strain versus cumulative number of cycles for WRC1 and SP1 specimens. 00.511.522.533.544.550200400600800100012001400160018002000 CYCLESAXIAL STRAIN (%) (%) SP1_C WRC1_C SP1_Reload WRC1_Reload Beginning of Re-Load A pplication After Rest Period Figure 8-7. Axial strain versus cumulative number of cycles. 153

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0510152025303540455055600200400600800100012001400160018002000CYCLESAXIAL/RADIAL ANGLE, (o) SP1_C WRC1_C SP1_Reload WRC1_Reload Beginning of Re-Load A pplication After Rest Period Figure 8-8. Angle versus cumulative number of cycles. The following figures display the axial and shear strain response during the loading and during the 8-hour rest versus time on logrithimic scale. 012345678910110100100010000100000Time (seconds)SHEAR STRAIN (%) WRC1 SP1 28% Drop 28% Drop F igure 8-9. Shear strain versus time 154

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00.511.522.533.544.55110100100010000100000Time (seconds)AXIAL STRAIN (%) (%) WRC1 SP1 23% Drop 29% Drop Figure 8-10. Axial stain versus time. The first observation prior to resting is that both the WRC1 and SP1 specimens behaved in a manner similar to the initial uninterrupted tests. At the end of the 8-hour rest period, both specimens appear to recover approximately 23 to 28% of the strain. Most of this strain is recovered after 15 minutes after the load was removed. Upon re-loading after the 8-hour rest period both the WRC1 and SP1 entered almost immediately into a failure/rupture response, lasting 20 to 50 cycles despite the recovery in strain. In the previous tests, the mixtures lasted 2 to 4 times longer in there respective tertiary ranges prior to failure. The almost immediate jump into failure upon load re-application is similar to the strain rates in the primary range. Load Relaxation During Secondary Phase and Re-Loading The third set of testing sought to stop the loading when the specimens were approximately mid-way through their respective Secondary Ranges 503 cycles for the SP1 mixture and 650 155

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cycles for the WRC1 mixture. Again, an 8-hour rest period was allowed prior to re-loading. Both mixtures regained/recovered approximately 6% of their radial and axial strain during the 8-hour rest period -less than one-quarter of the strain recovered when the tests was stopped in the tertiary range. The results are presented in the following figures. 02468101214160200400600800100012001400160018002000CYCLESSHEAR STRAIN (%) SP1 WRC1 SP1_Reload WRC1_Reload Figure 8-11. Shear strain versus cumulative number of cycles 156

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0123456789100200400600800100012001400160018002000CYCLESAXIAL STRAIN (%) SP1 WRC1 SP1_Reload WRC1_Reload Figure 8-12 Axial strain versus cumulative number of cycles. 0510152025303540455055600200400600800100012001400160018002000CYCLESANGLE OF AXIAL/RADIAL, (o) SP1 WRC1 SP1_Reload WRC1_Reload Figure 8-13. Angle versus cumulative number of cycles. 157

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Unlike the re-load application after resting during the tertiary range, the mixtures responded as if they had not been loaded before. The following figures display the results as if the re-loading was a separate test/specimen altogether beginning: 0123456789100200400600800100012001400160018002000CYCLESSHEAR STRAIN (%) SP1_Reload WRC1_Reload Figure 8-14. Shear strain versus number of cycles 158

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00.511.522.533.544.550200400600800100012001400160018002000CYCLESAXIAL STRAIN (%) (%) SP1_Reload WRC1_Reload Figure 8-15. Axial strain versus number of cycles. 0510152025303540455055600200400600800100012001400160018002000CYCLESANGLE OF AXIAL/RADIAL, (o) SP1_Reload WRC1_Reload Figure 8-16. Angle versus number of cycles. 159

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Both mixtures responded initially in a similar manner prior to load relaxation as previously for the separate mixtures. After the load was re-applied after the 8-hour rest period, both mixtures responded in a very similar manner in terms of both shear and axial strain rate. The SP1 mixture had been stopped at 500 cycles. It then was re-loaded and failed after an additional 1,050 cycles (1,550 cumulative cycles). The WRC1 mixture was stopped at 650 cycles and failed at 1,050 cycles (1,700 cumulative cycles). Both mixtures enter their respective tertiary ranges at the same radial and axial strains (for the re-load test, not cumulative) as when there was no load stoppage. Evaluations The hypothesis of this research centered on testing mixtures at high shear under low confinements -the stress state believed to result in instability rutting. In addition, the elevated temperature ranges and relatively un-aged pavements that typically undergo instability rutting suggest that resistance to instability rutting may be tied to a mixtures aggregate make-up. The preliminary testing results indicated that the WRC1 withstood the loading conditions better than the SP1 mixture took more load applications to result in failure. It appeared that the radial strain response of both mixtures was similar. However, the SP1 mixture did have more initial shear strain during its primary response phase. When both mixes reached between 2.5% and 3.0% axial strain they entered into their respective tertiary ranges, eventually leading to failure. The SP1 Mixture entered this range at approximately 900 to 1,000 cycles while the WRC1 mixture at about 1,200 cycles. This is more apparent when observing the angle response over the duration of the test the SP1 mixture has a high angle meaning a higher proportion of its strain was axially than the WRC1 mixture. It was this axial strain that contributed to a specimen reaching failure. 160

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Next, the phenomenon of healing was investigated through a series of two tests. Each mixture had a similar response when re-loaded after being stopped during their respective tertiary phases each went into immediate failure. This suggests that even with the large recoverable strain, the specimens of both mixtures had achieved a state of irrecoverable damage. It is likely, that the tertiary state is one in which the specimens had achieved the formation of an irrecoverable slip or failure plane within the aggregate orientation -thus the inability of both mixtures upon re-loading to absorb additional loading cycles -therefore no healing and failure upon re-loading. This further suggests that because of the large recovered strain, that a threshold in terms of total strain may only be applicable during initial loading, and not during re-loadings. However, there was a little difference in the shear and axial strain response between the SP1 and WRC1 mixtures when re-loading occurred after stopping during the secondary range. Prior to load relaxation, the two mixtures were behaving as they had typically been both mixtures had a similar radial strain response, the SP1 mixture has slightly higher axial strain rate. Both tests were then stopped mid-way through their respective secondary response ranges and allowed the rest under isotropic relaxed stated conditions for 8 hours. When the load re-applied, the response between the two mixtures was almost identical in terms of their axial and radial strain response. The SP1 mixture continued for an additional 1,050 cycles (a cumulative 1,550 cycles) and the WRC1 continued for an additional 1,050 cycles (a cumulative 1,700 cycles). Both mixtures entered their respective tertiary ranges at approximately 2.5% axial strain when viewing the re-loading as a separate test, not as a cumulative test. Thus, if the specimens had no memory, the two mixtures performed almost the same and the differences between them are almost impossible to discern which mixture is the better performer. 161

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Conclusions The hollow cylinder device is a testing apparatus that is useful for imposing a variety of stress states upon asphalt specimens. This research centered on testing asphalt mixtures under high shear at low confinement and at elevated temperature to simulate the critical conditions which are believed to result in instability rutting. Two mixtures when tested under these conditions performed in a manner consistent with their known resistance to instability rutting. A mixturess ability to resist axial strain under the critical laboratory imposed stress conditions appeared to be was the determining factor in reaching failure. The tests also indicate that some mixtures may improve their performance relative to other mixtures when allowed to heal during the secondary response phase. Further research is necessary to evaluate and determine the nature of a specific that allows it to resist instability rutting better than other mixtures. The hollow cylinder devise created here at the University of Florida will help to investigate these issues due to its ability to impose a variety of stress states on specimens not allowed by other testing devices. 162

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CHAPTER 9 CONCLUSION Summary of Work Instability rutting is a complex phenomenon seen in asphalt pavements. Instability rutting is rutting which is confined only to the asphalt concrete layer, typically in the top 2 to 3 inches, and generally occurs during the first hot summer of the pavements life. Instability rutting in asphalt pavements is primarily due to the lateral displacement of material within the asphalt concrete layer. Three-dimensional finite element modeling of actual tire contact stresses identified what are believed to be the critical stress states the result in the propagation of instability rutting high shear at low confinement. These stress conditions are not identified by traditional representation of radial tire loads as uniform vertical loads. Testing under high shear at low confinement conditions would lead to a better understanding of the mechanisms behind instability rutting. A testing device was therefore created that could impose these critical stress states upon asphalt specimens --the hollow cylinder testing device (HCTD) at the University of Florida. Preliminary testing indicated that a good performing mixture might be the result of its ability to resist axial strain (dilation and/or expansion) better than poor performing mixtures under critical stress states. A HCTD is a complex device that is primarily a research tool. The HCTD is not a screening tool to predict instability rutting in mixtures that can enjoy widespread use nor is the protocol for such a device ready base don the limited studies conducted to date. However, prior to development of a simple screening device, an understanding of the mechanisms behind instability rutting is necessary. This is such that any future device, protocol, or methodology that 163

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is developed will be based on a sound mechanistic basis rather than an empirical one which may or may not be capturing the essence of the mechanism behind instability rutting. Future Work The HCTD is a versatile laboratory tool. The creation of the HCTD will allow researchers the ability to tests under a variety of stress states that cannot be obtained by many current devices. Based on the results of the preliminary findings the following topics should be further investigated: The influence on aggregate gradation on the ability of a mixture to resist instability rutting under critical stress states imposed by the HCTD. The influence on gradation and aggregate arrangement during testing under critical stress conditions to determine a failure limit. The incorporation of healing and failure limits to evaluate a mixtures potential for instability rutting. The verification of material response by simpler tests through confirmation by more complex stress states in the HCTD. The use of the HCTD through internal pressure greater than the external pressure induce hoop stress -in the study of cracking. 164

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LIST OF REFERENCES Alarcon, A., J.L. Chameau, and G.A. Leonards, A New Apparatus for Investigating the Stress-Strain Characteristics of Sands, Geotechnical Testing Journal, Vol. 9(4), 1986, pp. 204-212. Alavi, S.H. and C.L. Monismith, Time and Temperature Dependent Properties of Asphalt Concrete Mixes Tested as Hollow Cylinders and Subjected to Dynamic Axial and Shear Loads, Journal of the Association of Asphalt Paving Technologists, Vol. 63, 1994, pp.152-181. Alhborn, G., Elastic Layered System with Normal Loads, Berkeley, California: ITTE, University of California, 1972. Ali, H.A. and S.D. Tayabji, Evaluation of Mechanistic-Empirical performance Prediction Models for Flexible Pavements Transportation Research Record No. 1629, Transportation Research Board, Washington D.C., 1998, pp.169-180. Arnold, G., D. Hughes, A.R. Dawson, and D. Robinson, Design of Granular Pavements, Proceedings of the 8th International Conference on Low-Volume Roads, 2003. Bahia H.U., T. Friemel, P. Peterson, and J. Russell, Optimization of Constructability and Resistance to Traffic: A New Design Approach for HMA Using the Superpave Gyratory Compactor, Journal of Association of Asphalt Paving Technologists, Vol. 67, 1998, pp.189-203. Bathe, K., ADINA System 7.5, Users Manual, ADINA R&D, Inc., Watertown, MA, 2001. Bode, G., Forces and Movements Under Rolling Truck Tires, Automobiltechnische Zeitschrift, Vol. 64, No. 10, 1962, pp. 300-306. Boulbibane, M., D. Weichert, and L Raad, Numerical Application of Shakedown Theory to Pavements with Anisotropic Layer Properties, Transportation Research Record No.1687, Transportation Research Board, Washington D.C., 1999, pp.75-81. Brown, S.F., Laboratory Testing for Use in the Prediction of Rutting in Asphalt Pavements, Transportation Research Record No. 616, Transportation Research Board, Washington D.C., 1976, pp.22-27. Brown, S.F. and K.E. Cooper, A Fundamental Study of the Stress-Strain Characteristics of Bituminous Materials, Journal of the Association of Asphalt Paving Technologists, Vol. 48, 1990, pp. 476-498. Buttlar, W.G., G. Al-Khateeb, and D. Bozkurt, Development of a Hollow Cylinder Tensile Tester to Obtain Mechanical Properties of Bituminous Paving Mixtures, Journal of the Association of Asphalt Paving Technologists, Vol. 68, 1998, pp. 369-403. 165

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BIOGRAPHICAL SKETCH Marc Edmond Novak was born in Rahway, New Jersey on November 14, 1975 to John and Elaine Novak. He attended Jesuit High School in Tampa, FL and after graduation attended the University of Florida. Marc Novak graduated with his Bachelor of Science degree in Civil Engineering from the University of Florida in 1999 and then joined the graduate program. He obtained his Master of Engineering in 2000. Marc Novak was selected as an Alumni Research Fellow to pursue a doctoral degree in the field of pavement engineering at the University of Florida. In the fall of 1996, Marc was initiated into Tau Beta Pi where he met his future wife, Miss Patricia Kwong. In February of 2003, Marc married Patricia at St Lawrence Catholic Church in Tampa, Florida. In 2004, he began to work in Tampa for a materials and geotechnical engineering consulting firm. Upon completion of his doctoral program, Marc intends to continue in the consulting field, enjoying time with his wife, and eventually pursue a career in research and academia. 176