INVESTIGATING ASPECTS OF AGGREAGTE PROPERTIES

THAT INFLUENCE ASPHALT MIXTURE PERFORMANCE

by

Clint Miller

Kamilla L Vasconcelos

Dallas N Little

and

Amit Bhasin

Research Report For DTFH61-06-C-00021

Texas A & M University, College Station

and

The University of Texas at Austin, Austin

June 2011

DISCLAIMER

The contents of this report reflect the views of the authors, who are responsible for the facts

and the accuracy of the information presented herein. Mention of trade names or commer-

cial products does not constitute endorsement or recommendations for use.

ACKNOWLEDGMENTS

The authors recognize that support was provided by Federal Highway Administration through

the Aggregates Foundation for Technology, Research and Education and the National Sand

Stone and Gravel Association to the International Center for Aggregate Research.

v

ABSTRACT

The physical and physio-chemical properties of aggregates influence different aspects

of the performance of asphalt mixtures. Some of these properties are inherent to the nature

and source of the aggregate (e.g. mineral composition) while other properties are, at least to

some extent, influenced by the processes used to produce the asphalt mixture (e.g. mixing

and compaction temperature). This research was conducted in two parts.

The first part of this research was to evaluate the role of asphalt mixture production

temperatures on the performance of Warm Mix Asphalt (WMA). Results from the dynamic

mechanical analysis of fine aggregate matrix (FAM or fine aggregate-asphalt mortar) spec-

imens indicate that for a given material combination, shear modulus and fatigue cracking

resistance of the FAM typically decreased when the specimens were produced at a rela-

tively lower mixing and compaction temperature compared to standard hot mix. Results

from tests conducted using the micro calorimeter demonstrate that for the non-porous ag-

gregates used in this study, lower aggregate pretreatment temperatures (within the range

of 90oC to 150oC) did not significantly impact the adhesive bond strength with asphalt

binders.

The second part of this research was to evaluate the physio-chemical properties of com-

mon minerals found in aggregates that influence their interaction with asphalt binders.

Twenty-two minerals that commonly comprise the surfaces of aggregates were carefully

selected and tested using the Universal Sorption Device (USD) to determine their surface

free energy components. The researchers believe this to be the most extensive catalog

of mineral surface properties ever compiled. This catalog can serve as a reference for

properties that can be used to explain specific interactions between minerals and materials

interacting with these minerals including but not limited to bitumen and water.

The dimensionless energy ratio of adhesive bond strength between aggregate (mineral)

and bitumen in dry condition to the bond strength in the presence of water was calculated

for two typical asphalt binders. Results demonstrate a wide variation in this energy ra-

tio based on the mineral considered. An important finding from this study was that certain

minerals developed a thermodynamically favored bond with bitumen compared to the com-

peting bond with water. Previous testing on aggregates and computation of energy ratios

always revealed that aggregate surfaces thermodynamically favor a bond with water com-

pared to bitumen. However, aggregates are not generally purely comprised of one mineral

type, and the large majority of the minerals tested demonstrated a potential to strip. This

finding requires more evaluation.

vi

TABLE OF CONTENTS

List of Figures ix

List of Tables xi

Chapter 1. Introduction 1

1.1 Background and Objectives . . . . . . . . . . . . . . . . . . . . . . . . . . 1

1.2 Report Structure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1

Chapter 2. Background and Literature Review 3

2.1 Aggregate Properties Related to Strength and Stability . . . . . . . . . . . . 3

2.2 Aggregate Properties Relevant to Adhesion . . . . . . . . . . . . . . . . . 6

2.2.1 Influence of aggregate shape . . . . . . . . . . . . . . . . . . . . . 6

2.2.2 Influence of physio-chemical properties . . . . . . . . . . . . . . . 7

2.3 Impact of Construction Operations on Aggregate Properties Relevant to

Adhesion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.3.1 Moisture on Aggregate Surfaces During Production . . . . . . . . . 9

2.3.2 Aggregate Storage and Contamination . . . . . . . . . . . . . . . . 11

2.4 Test Methods to Measure Aggregate Properties Relevant to Adhesion . . . . 12

2.4.1 Universal Sorption Device . . . . . . . . . . . . . . . . . . . . . . 13

2.4.2 Wilhelmy Plate Method . . . . . . . . . . . . . . . . . . . . . . . . 14

2.4.3 Micro-calorimeter . . . . . . . . . . . . . . . . . . . . . . . . . . . 14

2.4.4 Aggregate Imaging System (AIMS) . . . . . . . . . . . . . . . . . 15

2.5 Test Methods to Measure Performance of Asphalt Composites . . . . . . . 15

2.6 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17

Chapter 3. Influence of Mixture Production Temperature on Adhesive Prop-

erties of Aggregates 19

3.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.2 Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

3.3 Influence of Production Temperature on Adhesion and Surface Properties

of Aggregates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

3.3.1 Overview of the test method . . . . . . . . . . . . . . . . . . . . . 21

vii

3.3.2 Sample preparation . . . . . . . . . . . . . . . . . . . . . . . . . . 23

3.3.3 Test and analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.3.4 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

3.3.5 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

3.4 Influence of Production Temperatures on Performance of FAM . . . . . . . 31

3.4.1 Sample preparation . . . . . . . . . . . . . . . . . . . . . . . . . . 31

3.4.2 Test and analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . 32

3.4.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

3.4.4 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34

Chapter 4. Physio-Chemical and Adhesive Properties of Common Minerals

in Aggregates 39

4.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39

4.2 Factors That Influence Surface Free Energy of Aggregates . . . . . . . . . 39

4.2.1 Non-polar active sites . . . . . . . . . . . . . . . . . . . . . . . . 39

4.2.2 Polar active sites . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

4.2.3 Coatings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41

4.3 Materials and Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42

4.3.1 Elemental analysis of the minerals . . . . . . . . . . . . . . . . . . 42

4.3.2 Pretreatment of minerals for surface energy testing . . . . . . . . . 43

4.3.3 Testing with the universal sorption device . . . . . . . . . . . . . . 43

4.4 Results from Surface Energy Testing . . . . . . . . . . . . . . . . . . . . . 45

4.4.1 Carbonates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

4.4.2 Sulfates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

4.4.3 Phyllosilicates . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47

4.4.4 Feldspars . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

4.4.5 Oxides . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 49

4.4.6 Nesosilicates and Inosilicates . . . . . . . . . . . . . . . . . . . . 50

4.5 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

Chapter 5. Concluding Remarks 63

5.1 Influence of Mixture Production Temperature . . . . . . . . . . . . . . . . 63

5.2 Physio-chemical and adhesive properties of mineral aggregates . . . . . . . 64

References 67

viii

LIST OF FIGURES

Figure 2.1. Typical temperature profile in a drum mix plant during hot mix

asphalt production. . . . . . . . . . . . . . . . . . . . . . . . . . 10

Figure 3.1. Schematic illustrating the working principle of the differential

micro-calorimeter (Bhasin 2006) with a reaction cell containing

fine aggregates and an empty reference cell; the syringes on cells

are filled with the liquid used to immerse the aggregate sample. . . 23

Figure 3.2. The micro-calorimeter (left) is shown with a typical vial contain-

ing an aggregate sample (right); the asphalt binder is in solution

form in the syringe ready to be injected into the vial with the ag-

gregate to measure the heat of immersion. . . . . . . . . . . . . . 25

Figure 3.3. Typical heat flow measured using the micro-calorimeter when a

solid (aggregate) is immersed in a probe liquid. . . . . . . . . . . 25

Figure 3.4. Influence of aggregate treatment temperature on the total energy

of immersion with asphalt binders. . . . . . . . . . . . . . . . . . 28

Figure 3.5. Influence of aggregate treatment temperature on the heat of im-

mersion for aggregate RK. . . . . . . . . . . . . . . . . . . . . . 28

Figure 3.6. Influence of aggregate treatment temperature on the heat of im-

mersion for aggregate RL. . . . . . . . . . . . . . . . . . . . . . 29

Figure 3.7. Influence of aggregate treatment temperature on the heat of im-

mersion for aggregate RD. . . . . . . . . . . . . . . . . . . . . . 29

Figure 3.8. Influence of aggregate treatment temperature on computed adhe-

sive bond strength with different asphalt binders. . . . . . . . . . . 30

Figure 3.9. A typical plot used to identify the number of load cycles to failure;

note that curve with the number of load cycles multiplied by the

normalized reduction in shear modulus has a well defined peak

that is identified with failure. . . . . . . . . . . . . . . . . . . . . 33

Figure 3.10. Influence of mixing and compaction temperature on the shear

modulus of dry specimens (average shown with +/- one standard

deviation). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

ix

Figure 3.11. Influence of mixing and compaction temperature on the fracture

resistance of dry specimens (average shown with +/- one standard

deviation). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

Figure 3.12. Influence of moisture conditioning on linear viscoelastic property

of different FAM specimens. . . . . . . . . . . . . . . . . . . . . 36

Figure 3.13. Influence of moisture conditioning on fracture resistance of dif-

ferent FAM specimens. . . . . . . . . . . . . . . . . . . . . . . . 36

Figure 4.1. Summary of total, polar and non-polar components of surface en-

ergies for minerals (arranged by group and in the order discussed

above). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51

Figure 4.2. Laboratory performance (tensile modulus) of mixtures compared

to ER2 * SSA (ER2 is a parameter based on surface free energies

of asphalt binder and aggregate and SSA is the specific surface

area of the aggregate). (After Little and Bhasin, 2006). . . . . . . 53

Figure 4.3. Laboratory performance (plastic deformation) of mixtures com-

pared to ER2 * SSA parameter (ER2 is a parameter based on sur-

face free energies of asphalt binder and aggregate and SSA is the

specific surface area of the aggregate) (After Little and Bhasin,

2006). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

Figure 4.4. Energy ratio representing inherent resistance to moisture induced

damage for mineral-asphalt binder pairs presented from lowest

to highest and for asphalts AAB and AAD (higher energy ratio

indicates better resistance to moisture damage). . . . . . . . . . . 55

Figure 4.5. Comparison of fatigue lives of moisture conditioned asphalt mix-

tures to dry asphalt mixtures for bitumen AAB and AAD. . . . . . 56

Figure 4.6. Surface energy trends within a mineral group and identification of

minerals that preferred a bond with bitumen rather than water. . . 56

Figure 4.7. Base surface energy versus moles of organic carbon on mineral

surfaces. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58

x

LIST OF TABLES

Table 4.1. List of Minerals for Surface Energy Measurements . . . . . . . . 59

Table 4.2. List of Minerals for Surface Energy Measurements . . . . . . . . 60

Table 4.3. Summary of Surface EnergyMeasurements ofMinerals (ergs/cm2)

61

xi

CHAPTER 1. INTRODUCTION

1.1 BACKGROUND AND OBJECTIVES

The physical and physio-chemical properties of aggregates influence different aspects of

the performance of asphalt mixtures. Some of these properties are inherent to the nature

and source of the aggregate (e.g. mineral composition) while other properties are, at least

to some extent, influenced by the processes used to produce the aggregate or the asphalt

mixture (e.g. mixing and compaction temperature). The importance of characterizing the

physical, chemical and thermodynamic properties of the aggregates that influence the per-

formance of the asphalt mixtures cannot be underestimated.

The broader objective of this project was to understand the importance of inherent

physio-chemical properties of the aggregate and factors related to aggregate and mixture

production that influence mixture performance. More specifically, this research evaluated

the following:

1. The role of asphalt mixture production temperatures on the performance of WMA.

2. The physio-chemical properties of common minerals found in aggregates and the

influence of these properties on the ability of the mineral to bond with asphalt binders.

1.2 REPORT STRUCTURE

Chapter 2 of this report presents a comprehensive literature review on the aggregate prop-

erties that influence the strength and stability of asphalt mixtures as well as aggregate prop-

erties that influence its adhesion to asphalt binders and resistance to moisture induced dam-

age. This chapter also presents a discussion on the impact of production processes that

influence these aggregate properties, a summary of test methods that are available to char-

acterize the aggregate properties relevant to adhesion, and a summary of test methods that

are used to characterize the performance of asphalt mixtures.

Chapter 3 presents the findings from the tests conducted to evaluate the influence of

mixture production temperatures on the performance of asphalt mixtures. The first part

of this chapter evaluates the influence of mixture production temperatures on the physio-

chemical or surface properties of aggregates relevant to adhesion. The second part of this

chapter evaluates the influence of mixture production temperatures on the cracking and

1

moisture damage resistance of fine aggregate matrix (FAM or asphalt binder - sand com-

posites) using a dynamic mechanical analyzer. The fine aggregates used in the FAM speci-

mens correspond to the fine aggregate from different full-graded mixtures.

Chapter 4 presents the findings from the tests conducted to evaluate the surface free en-

ergy of pure minerals commonly found in aggregates. Twenty two pure minerals were in-

cluded in this study and their surface properties were cataloged. This chapter also presents

analysis that demonstrates the inherent resistance to moisture induced damage offered by

different minerals. Finally, Chapter 5 presents concluding remarks and a summary of find-

ings from this study.

2

CHAPTER 2. BACKGROUND AND LITERATURE REVIEW

It is important to assimilate knowledge regarding the influence of aggregate properties that

have a significant impact on asphalt mix performance as well as the test and analytical

methods that can be used to determine these properties. A detailed review of applicable

literature was conducted to assemble information from domestic and international agencies

that have explored the influence of inherent aggregate characteristics (surface energy, shape

characteristics) on mixture performance.

2.1 AGGREGATE PROPERTIES RELATED TO STRENGTH AND STABILITY

The physical and chemical properties of aggregates are known to influence the performance

of asphalt mixes. For example, many studies emphasize the role of aggregate shape in

controlling the performance of asphalt mixtures, especially resistance to fatigue cracking

and rutting (Masad et al., 2005). These studies conducted experiments that focused on the

influence of fine aggregate, coarse aggregate, or the combined effect of fine and coarse

aggregate on the mechanical properties and performance of asphalt mixtures.

Campen and Smith (1948) found that when crushed fine aggregates were used instead

of natural rounded aggregates, the stability of dense-graded HMA mixtures increased from

30 to 190 percent. They measured stability using the bearing-index test. Ishai and Gellber

(1982) used the packing volume concept developed by Tons and Goetz (1968) to quantify

the geometric irregularities of a wide range of aggregate sizes. The HMA mixtures con-

taining different aggregates types were evaluated by Ishai and Gellber (1982) for Marshall

stability and flow, resilient modulus, and split tension strength. The results showed that

there was a significant increase in stability with an increase in the geometric irregularities

of the aggregates. No correlation was found between geometric irregularities and resilient

modulus or indirect tensile strength of the HMA mixtures.

Kalcheff and Tunnicliff (1982) evaluated the affect of fine aggregate shape on HMA

properties. HMA mixtures were tested using Marshall stability, repeated-load triaxial com-

pression, static indirect tensile strength, and repeated-load indirect tensile resistance tests.

They found that the use of manufactured sand instead of natural sand improved the mix-

tures resistance to permanent deformation from repeated loads, tensile strength, and tensile

fatigue resistance. Winford (1991) reached the same conclusion by relating mechanical

properties of HMA, such as those obtained from the static confined creep test, to the type

3

of fine aggregate in the mix.

Herrin and Goetz (1954) reported that when the amount of crushed gravel in the coarse

aggregate increased, the triaxial compressive strength of the dense-graded HMA was not

significantly influenced. However, the strength of the open-graded HMA mixture increased

significantly when the percentage of angular coarse aggregates was increased. Field (1958)

found a considerable increase in HMA Marshall stability due to an increase in the percent-

age of crushed coarse particles. The influence of crushed gravel coarse aggregate on the

properties of dense-graded HMA mixtures was also investigated by Kandhal and Wenger

(1975). They found that the Marshall stability of a dense-graded mix decreased with an

increase in uncrushed gravel particles. However, the differences among the mixes were not

significant. They also noted that there was no significant difference in the tensile strength

of HMA mixtures containing crushed and uncrushed coarse aggregates.

Sanders and Dukatz (1992) reported on the influence of coarse aggregate angularity on

permanent deformation of four interstate sections of HMA pavements in Indiana. One of

the four sections developed permanent rutting within two years of service. They found that

HMA mixtures used in the base course and the surface course of the rutted section had

lower amounts of angular coarse aggregate compared to the other three sections.

Kandhal et al. (1998) pointed out that only a few studies had been conducted to examine

the influence of flat and elongated coarse aggregate particles on HMA strength compared

to the number of studies that addressed coarse aggregate angularity. The presence of ex-

cessive flat and elongated aggregate particles is undesirable in HMA mixture because such

particles tend to break down (especially in open-graded mixtures) during production and

construction, thus affecting the durability of HMA mixtures.

A study by Li and Kett (1967) found that the dimension ratio (width to thickness or

length to width) had no effect on Marshall or Hveem stability as long as the dimension

ratios were less than 3:1. The permissible percentage of flat and/or elongated particles (di-

mension ratio exceeding 3:1), that did not adversely affect the mix stability was determined

to be 30 percent or as much as 40 percent. Stephens and Sinha (1978) reported that HMA

mixes containing 30 percent or more flat particles (longest axis to shortest axis is more

than or equal to three) maintained higher void contents compared to some other blends

with lower percentages of flat particles. These mixes were compacted with equivalent ef-

forts using a kneading compactor.

Some studies focused on comparing the relative influence of fine aggregate and coarse

aggregate shape properties on the mechanical properties and performance of HMA. Lefebure

4

(1957) utilized the Marshall test to measure the stability of HMA mixtures containing a

crushed cubical coarse aggregate or crushed aggregates with flat and long particles com-

bined with natural sand or crushed sand. His study concluded that fine aggregate was the

most critical component of the HMA mixture. Its quantity and characteristics controlled, to

a large extent, the Marshall stability. Wedding and Gaynor (1961) evaluated the influence

of crushed coarse and fine aggregate on the Marshall stability of dense-graded HMA mix-

tures. Using crushed coarse aggregates caused a significant increase in stability compared

to uncrushed coarse aggregates. The use of crushed fine aggregates increased the stability

of mixes containing uncrushed coarse aggregates. However, the use of crushed fine ag-

gregates had a minimal effect on HMA stability when the mixes contained crushed coarse

aggregates.

Foster (1970) measured the resistance of dense-graded HMA mixtures to traffic by

using test sections. He concluded that HMA mixtures containing crushed coarse aggregate

showed no better performance than a mix containing uncrushed aggregates. The study

attributed this finding to the crushed fine aggregate, which controlled the capacity of the

mix to resist stresses induced by traffic.

The influence of shape, size, and surface texture of aggregate on stiffness and fatigue

response of HMA mixture was investigated and summarized by Monismith (1970). He

indicated that aggregate characteristics affect both stiffness and fatigue response of HMA

mixtures. Monismith (1970) recommended utilizing rough-textured materials with dense

gradation for thick pavements in order to increase mix stiffness and fatigue life; whereas,

it might be acceptable to utilize smooth-textured aggregates in thin pavements since they

produce less stiff mixtures resulting in increased fatigue life. Barksdale et al. (1992) eval-

uated the effect of aggregate on rutting and fatigue of HMA mixtures. Aggregate shape

was measured using image analysis techniques and the packing test developed by Ishai and

Gellber (1982) . They found that aggregate shape properties obtained from the packing test

were statistically related to the rutting behavior of selected HMA mixtures. A comprehen-

sive study by Kandhal et al. (1991) evaluated the factors that contribute to asphalt pavement

performance. They found that mixtures with less than 20 percent natural sand in the fine

aggregate had better performance than mixtures with more than 20 percent natural sand.

They also recommended using coarse aggregate having at least 85 percent of particles with

two or more fractured faces for heavy-duty wearing and binder courses.

A study conducted at the Texas Transportation Institute related an imaging index of

aggregate texture (fractal dimension) to the creep behavior of asphalt mixes (Yeggoni et al.,

5

1996). In this study, seven different aggregate blends of the same gradation but with varying

amounts of crushed coarse aggregate particles were prepared.

2.2 AGGREGATE PROPERTIES RELEVANT TO ADHESION

Geometric, chemical and thermodynamic properties of the aggregate surfaces are impor-

tant in determining the nature and durability of the bond with bitumen in dry condition

and in the presence of water. Loss of bond between the aggregate and the bitumen can

occur in two modes. The first mode is adhesive failure between the aggregate and the bi-

tumen, and the second mode is cohesive failure of a weak boundary layer that is formed

at the bitumen-aggregate interface. Several theories in the literature explain adhesive fail-

ure based on mechanisms related to mechanical bonding, thermodynamic interactions or

physical bonding, and chemical bonding. It is unlikely that any of these mechanisms or

theories can exclusively explain the quality of the bond between an aggregate surface and

bitumen. The relative influence of these mechanisms on the performance of any combina-

tion of asphalt and aggregate depends on the physical and chemical characteristics of both

the aggregate and the binder.

2.2.1 Influence of aggregate shape

Aggregate shape characteristics influence mechanical adhesion between the aggregate and

binder. Pocius (1997) described that three factors that contribute to mechanical adhesion:

physical “lock and key,” redistribution of stresses, and increased surface area. An important

aspect of resistance to moisture damage is the bond strength per unit area within the binder

and at the binder-aggregate interface. As the level of aggregate texture and angularity in-

creases, aggregate surface area increases, and the result is an increased total bond energy in

the mix. Tarrer and Wagh (1991) emphasize the importance of surface texture in mechan-

ical bonding of the bitumen and aggregate. They characterized macroscopic roughness of

aggregates in terms of its surface texture. Surface texture of aggregates at the microscopic

level can be significantly different from the surface texture at the macroscopic level. A

measure of the surface texture of aggregates at the microscopic level is the specific surface

areas of the aggregates. Bhasin and Little (2006) demonstrate that aggregates have a very

broad range of specific surface areas that varies from 0.1 to 10 m2/gm for the same size

fraction (passing 4.7 mm sieve and retained on 2.36 mm sieve). Specific surface areas of

aggregates were measured during the process of estimating their surface free energies using

6

the universal sorption device.

Hicks (1991) built upon the work by McBain and Hopkins and reported that it is gen-

erally accepted that aggregates with a porous, slightly rough surface promote adhesion by

providing a mechanical interlocking effect (McBain and Hopkins, 1929, 1925). A possible

adverse effect of angularity on moisture damage is an increased probability of puncturing

the asphalt film and allowing intrusion of water to the asphalt-aggregate surface.

The relationship between aggregate shape properties and HMA performance under wet

conditions was described by Masad et al. (2004). Researchers tested and analyzed HMA

mixtures with limestone, siliceous gravel, and granite aggregates. The granite aggregate ex-

hibited the highest levels of texture and angularity, followed by limestone, and then gravel

(Masad et al., 2004). The performance testing results indicated that the granite mix was

superior to the limestone mix under wet conditions. Researchers theorized that the granite

retained a superior level of bonding in the presence of water compared to the limestone

aggregate because of its higher texture and angularity.

2.2.2 Influence of physio-chemical properties

Thermodynamic properties such as surface free energies of the aggregate and bitumen are

used to explain the adhesion between these materials and the propensity of water to displace

bitumen from the bitumen-aggregate interface (Majidzadra and Brovold, 1968; Kanitpong,

2004). Curtis et al. (1993) determined the Gibb’s free energy of aggregates from adsorption

isotherms. Cheng (2002) demonstrated the use of Wilhelmy plate method and the Universal

Sorption Device (USD) to measure surface free energies of aggregate and bitumen. He

uses the surface free energies of these materials to calculate their interfacial bond strength

in wet and dry conditions based on thermodynamic relationships. Kim et al. (2004) and

Bhasin et al. (2006) demonstrate the correlation of these thermodynamic parameters to the

moisture sensitivity of an asphalt mix.

The chemical bonding theory suggests that the adhesion between the aggregate and

bitumen is due to the formation of weak chemical bonds between various polar functional

groups in the bitumen with the active minerals on the aggregate surface. Petersen (1984)

reported eight different types of functional groups that are typically found in bitumen and

can chemically interact and form bonds with the aggregate surface. Examples of weak

acid type functional groups are carboxylic acids and anhydrides, and examples of weak

base type functional groups are sulfoxides and pyridines. Jamieson et al. (1995a) report

that, chemical sites on aggregate surfaces associated with high affinity for bitumen include

7

elements such as aluminum, iron, magnesium and calcium. Elements associated with low

bonding affinity include sodium and potassium. Little et al. (2002) reported the formation

of calcium or sodium salts when carboxylic acid from the bitumen chemically interacts with

the available calcium or sodium on the aggregate surface. While sodium salts are readily

soluble in water and accelerate moisture damage, calcium salts are relatively insoluble and

therefore are more resistant to water induced damage. The utility of active fillers such

as hydrated lime and liquid anti-strip additives to retard moisture induced damage is also

explained based on such chemical interactions. Details of specific interactions between

various types of aggregates and bitumen functional groups are available in the literature.

In another study, a micro-calorimeter was used to quantify the effect of chemical and

physical interactions between model compounds (that represent bitumen functional groups)

and active minerals (Ensley and Scholz, 1972; Ensley, 1973). Evidence in the literature

supports the hypothesis that the amount of insoluble material produced at the interface, and

hence the resistance of the bond to moisture damage, depends on the amounts of chemical

reactants present on the aggregate surface and the bitumen (Institute, 2001). Methods such

as quantitative X Ray diffraction can be used to quantitatively determine the percentage

of active minerals on aggregate surfaces. This can be combined with the knowledge of

specific interactions between active minerals and functional groups from bitumen to predict

the behavior of bonding between the aggregate and the bitumen.

Another mode of loss of the aggregate-bitumen bonding is the failure in cohesion of

a weak boundary layer. Phenomenological observation indicates that when absorption of

bitumen on a porous aggregate occurs, the bitumen on the outside becomes hard and brittle

(Jeon and Curtis, 1990). This forms a weak boundary layer that makes the bond between the

bitumen and aggregate susceptible to premature damage. Podoll and Becker (1991) used

Surface Analysis by Laser Ionization (SALI) to examine water bonded surface of bitumen

and aggregate pairs. This technique essentially produces an elemental depth profile through

mass spectrometry. Calculation of relative abundances of various elements on the stripped

and unstripped areas indicates that the locations of failure of some bitumen-aggregate pairs

occurs on the aggregate side. Weak boundary layers of this nature may be intrinsic to the

aggregate used, or can be developed by dissolution of surface complexes (Jeon and Curtis,

1990) or the mineral itself in the presence of water. The latter is dependent upon the pH

of the contacting water. Cohesive failure due to dissolution of surface layers in carbonates

occur at pH levels lower than 6, while dissolution of silica minerals occurs at pH values

grater than about 8 (Jamieson et al., 1995b).

8

2.3 IMPACT OF CONSTRUCTION OPERATIONS ON AGGREGATE PROPER-

TIES RELEVANT TO ADHESION

2.3.1 Moisture on Aggregate Surfaces During Production

A number of factors during the production of asphalt mixtures influence the quality of the

adhesive bond between the aggregate and the asphalt binder. The mechanical forces during

mix production change the physical and chemical characteristics of the aggregate. The

physical changes are evident in aggregate abrasion and breakage that can change aggregate

shape characteristics and gradation. These changes, although not significant, affect the

mechanical bond between aggregates and the binder.

The temperature profile and the operation of a drum mix plant affects the aggregate

moisture content. A typical temperature profile along the depth of the drum is shown in

Figure 2.1. Aggregate movement along the drum is caused by gravity and the lifting flights

as the drum rotates. The flow of aggregates is retarded in the middle of the drum to allow for

better heat transfer from the exhaust gases so that the drying and heating of the aggregate

can take place. This can be accomplished by inserting a ring inside the drum, installing

reverse-angle flights to intercept aggregates, reducing the drum diameter in the middle, or

lowering the slope of the drum (HMA Handbook, 1991).

Although the temperature inside the drum decreases downstream, the aggregate tem-

perature does not follow the same profile. When aggregate is introduced into the drum,

its temperature increases as it moves downstream until it reaches a point before the drum

midpoint. After this point, the aggregate temperature remains almost constant as water

evaporates. The amount of time for which aggregate temperature remains constant de-

pends on moisture content and aggregate porosity. More porous aggregates may require

more time to remove moisture. Once most of the moisture has been removed from the ag-

gregate, its temperature begins to rise again. The moisture content of the mix at discharge

is almost always less than 0.5 percent.

Aggregate moisture content can have significant effect on aggregate coating and perfor-

mance. At the point where asphalt is introduced in the drum, the small amount of moisture

remaining on the surface of the aggregate causes the volume of the binder to expand by

foaming and helps to coat the aggregate. Therefore, when the average moisture content

of the incoming aggregate is very low, incomplete coating of the aggregate may occur.

On the other hand, excessive moisture content in the aggregate would lead to performance

problems as the trapped moisture will damage the adhesive binder-aggregate bond lead-

9

Figure 2.1. Typical temperature profile in a drum mix plant during hot mix asphalt

production.

ing to moisture related damage that can be manifested in loss of binder, disintegration of

aggregate, and premature fatigue cracking and rutting.

Parker and West (1992) evaluated the impact of residual moisture in aggregates on the

stripping potential of asphalt mixtures. The moisture content of the aggregates just prior to

mixing with the asphalt binder in a drum mix plant depends on the ambient temperature and

the moisture content of the stock piles. Parker and West reported that aggregates may not

be effectively dried by the rapid heating in drum dryers. They also reported that siliceous

aggregates were more sensitive to the presence of residual moisture compared to limestone

aggregates. In another study, Fwa and Ong (1994) reported that smaller aggregate par-

ticles had more moisture content expressed as a percentage of the dry aggregate weight

as compared to larger aggregate particles. This was attributed to the larger surface areas

per unit mass of the smaller aggregate particles. However, these authors also reported that

the time required for fully saturated aggregate particles to oven dry at 110°C was longer

for larger aggregate particles compared to smaller aggregate particles. The authors also

cite Lottman’s (1961) observation that moisture in deep pores of large aggregates may not

completely vaporize if the aggregates are not heated for sufficient duration of time.

10

Typically, a very small amount of surface adsorbed water or residual moisture is re-

quired to reduce the surface free energy of the aggregate surface (Adamson and Gast,

1997). For example, Legrand (1998) reports that as the temperature for preconditioning

certain types of silica is decreased, the percentage of silanol groups on the surface increases

and the total surface free energy decreases. Reduced surface free energy of the aggregate

translates into reduced adhesive bond strength with the asphalt binder.

In summary, existing literature demonstrates that the temperature and duration for which

aggregates are heated prior to mixing with the asphalt binder are important factors that dic-

tate the durability of asphalt mixtures. This is especially an important consideration for

warm asphalt mixtures that are produced at temperatures that are typically 20 to 30°C

lower than conventional mixtures. Although the workability of asphalt binders at lower

temperatures is achieved with the use of specialized additives, the qualitative and quan-

titative impact of residual moisture on aggregate surface on the durability of the asphalt

mixture remains to be addressed.

2.3.2 Aggregate Storage and Contamination

Aggregate storage is another factor that influences aggregate surface properties. Solid sur-

faces with high surface energies, such as aggregates, rarely have clean surfaces in their

native environments (Adamson and Gast, 1997). Freshly crushed aggregate surfaces in a

quarry are rapidly contaminated upon coming into contact with the atmosphere. Dust parti-

cles, organic materials, and water from atmospheric humidity adhere to the freshly crushed

surfaces. Adsorption of impurities on the aggregate surfaces continues while the aggregates

are stored in stock piles before being used in the hot mix. Even a single layer of organic

molecules can contaminate and alter the surface properties of the aggregates. Sources of

such contaminants include exhaust from plant mix operations in the vicinity of the stock

piles and atmospheric water. The presence of dust and organic substance on aggregate sur-

face affect its surface energy, and consequently the efficiency of binder coating and mix

performance.

Based on the testing of a limited number of aggregates, Little and Bhasin (2006) re-

ported that in most cases surface energy properties of freshly crushed aggregates did not

vary significantly from that of aged aggregates. However, a more detailed investigation of

the affect of aging on the surface properties of aggregates is required.

Johnson and Freeman (2002) report the presence of dust on aggregates as one of the

causes for its poor adhesion to asphalt binder. Tarrer and Wagh (1991) report that the pres-

11

ence of excessive dust and clay coatings on the aggregate surface inhibits its adhesion with

the asphalt binder and also provides a pathway for the water to get to the binder-aggregate

interface. Kandhal (1992) report that dust coated aggregates promote hydraulic scouring

and deterioration in the asphalt mixture. This report also cites examples of cases where

removal of dust from aggregate surface by cleaning improved the adhesion characteristics

of the coarse aggregate.

In summary, existing literature demonstrates that the contamination of aggregates due

to presence of dust on the aggregate surface and/or aging of the aggregate may affect the

quality of its adhesion with asphalt binder. However, very little information is available

that quantifies the impact of these factors on the adhesion with asphalt binder. This must

be considered for future work on aggregate properties.

2.4 TEST METHODS TO MEASURE AGGREGATE PROPERTIES RELEVANT

TO ADHESION

Adhesion is the tendency of dissimilar molecules to maintain intimate contact. Adhesion

involving geologic materials is controlled by the interfacial characteristics of rocks and

minerals, specifically the surface energy of natural minerals and rocks. Surface energy is a

thermodynamic construct defined as the work necessary to form a unit area of surface by a

process of division (Parks, 1990; Shuttleworth, 1950). Conceptually surface energy can be

considered as the amount of energy lost when the molecular bonds that are normally filled

inside a solid remain unfilled as a result of being at the edge of the solid (van Oss et al.,

1988a).

Surface energy of natural substances can be divided into a polar and non-polar com-

ponent. The non-polar (also referred to as the van der Waals or dispersive component) is

present in all molecules to varying degrees. These forces are produced by dipole-dipole in-

teractions and induced dipole interactions. Van der Waals interactions are generally weaker

than electrostatic interactions; however their influence extends over longer distances (van

Oss et al., 2001). Polar or non-dispersive forces are found where electron donor/electron

acceptor interactions take place and are further divided into electron donor (Lewis base

like), and electron acceptor (Lewis acid like) components (van Oss et al., 1988b,a, 2001).

The polar or Lewis acid-base interactions are mainly hydrogen donor and hydrogen accep-

tor reactions (Fowkes, 1966). However, it is useful to define “polar” more broadly for all

electron acceptor-donor interactions.

Using Good-van Oss Theory (vOGT) the total surface energy can be found by combin-

12

ing these elements as follows:

γTotal = γLW +2√

γ+γ− (2.1)

In equation 2.1, γTotal is the total surface energy, γLW is the Lifshitz–van der Waals or

dispersive component, γ+ is the Lewis acid component, and γ− is the Lewis base compo-

nent (van Oss et al., 1988b,a).

Bhasin and Little (2006) and Hefer et al. (2006) used the Universal Sorption Device and

the Wilhelmy plate device to determine the surface free energy components of aggregates

and asphalt binders, respectively. They demonstrated a good correlation between parame-

ters derived from these surface energy measurements to field and laboratory performance

of asphalt mixtures.

2.4.1 Universal Sorption Device

The universal sorption device (USD) is used to determine the surface energy components

of the aggregates. Tests with non-polar probe vapors such as nHexane are also used to

determine the specific surface areas of different aggregates (Bhasin, 2006). Most vapor

sorption methods described in the literature (Bilinski and Holysz, 1999; Chen and Dural,

2002) use very finely divided solids and a sample mass of about 1gm or less. These methods

cannot be effectively employed to measure surface free energies of aggregates on account

of their heterogeneity and difficulties in obtaining a representative aggregate sample. One

of the capabilities of the USD is its ability to test larger sample sizes of about 25 grams

which enables testing of representative aggregate samples. The aggregate size fraction

typically used in this device is between ASTM #4 and ASTM #8 sieves. Since surface

energy of an aggregate is an intrinsic material property it is independent of its geometry

and it is not required to test every size fraction of the same aggregate. This is true unless

the size fraction is extremely fine where the general crystal structure of the surface is altered

affecting its surface properties. Such fines are more likely to be considered as part of the

bitumen mastic rather than an aggregate bound by the bitumen. It must also be noted that

the size fraction tested with the USD have the same mineralogy as the coarse and fine

aggregates used in a typical full graded asphalt mixture. In some cases, the aggregate used

in an asphalt mixture may be a blend of aggregates representing two different sources. In

such cases, samples representing each source are tested separately.

13

2.4.2 Wilhelmy Plate Method

The Wilhelmy plate method is used to determine the surface free energy components of the

asphalt binder. Although this method does not measure aggregate properties, the surface

energy components of the binder are required as an input in order to quantify its work of

adhesion with an aggregate in dry condition and the release of free energy when water

displaces the binder from a binder-aggregate interface (Cheng, 2002; Hefer et al., 2006).

Therefore, it is important to measure the surface free energy of typical asphalt binders for

further analysis. Bhasin et al. (2007) used the Wilhelmy plate device to demonstrate the

impact of modification processes such as aging of asphalt binder, addition of liquid anti

strip agent, and addition of polymers on the work of adhesion using different aggregates.

2.4.3 Micro-calorimeter

The micro-calorimeter can be used to measure the total energy of a solid-liquid or a liquid-

liquid interaction in the form of enthalpy of immersion. Ensley et al. (1984) determined

the heat of immersion of different aggregates in different asphalt binders using a micro-

calorimeter. He related the measured heat of immersion to the rutting performance and

moisture sensitivity of asphalt mixes. Bhasin and Little (2006) employed the micro-calorimeter

using different approaches to determine properties of aggregates relevant to adhesion with

asphalt binders. They used the micro-calorimeter to measure the enthalpy of immersion of

aggregates in different pure homogenous probe liquids in order to compute the total sur-

face energy components of the former. Similarly, the enthalpy of immersion of aggregates

in water (at room temperature) and asphalt binder (at mixing temperature) was used as a

measure of hydrophilicity of the aggregate and the total energy of adhesion between the

aggregate and the binder, respectively. The enthalpy of immersion of aggregates in asphalt

binder can also be measured at room temperature by using the asphalt binder in a suitable

solvent. One of the advantages of using a micro-calorimeter is that it is a relatively faster

technique compared to the USD. However, the specific surface area of the aggregates must

be known a-priori in order to provide a general interpretation of the test results. Another

advantage of the micro-calorimeter is that the aggregate surface can be pre treated to rep-

resent a variety of different conditions prior to the test. This is in contrast to the USD that

requires the test to be conducted always from a clean dry aggregate surface conditioned

under vacuum. This advantage of the micro-calorimeter allows it to be used for surface

characterization of aggregates used in the production of warm asphalt mixtures.

14

2.4.4 Aggregate Imaging System (AIMS)

Masad et al. (2005) report the use of an Aggregate Imaging System (AIMS) to rapidly

measure the shape, form, and texture characteristics of aggregates. This system is equipped

with top lighting, back lighting, and autofocus microscope. These are necessary elements

of the system to capture black and white images, and gray scales images at different mag-

nifications for the analysis of all shape characteristics. The system is computer-controlled

to achieve motion in the x, y, and z directions as well as magnification. Masad et al. (2005)

developed a system for classification of aggregates based on their shape characteristics.

The AIMS was used in the preliminary portions of this study to determine the shape char-

acteristics of the aggregates.

2.5 TEST METHODS TO MEASURE PERFORMANCE OF ASPHALT COMPOS-

ITES

The test methods reviewed in the previous section are useful to determine properties of

aggregates that influence its adhesion with the asphalt binder. Mechanical tests and perfor-

mance modeling of fine aggregate matrix or whole asphalt mixture is necessary in order to

quantify the impact of change in these properties on the performance of the asphalt mixture.

The fine aggregate matrix (FAM) represents the portion of the asphalt mixture that

comprises of the asphalt binder, fines (material passing #200 sieve), and fine aggregates

(material passing #16 sieve). The FAM is an important and critical component of the

asphalt mixture that holds the coarse aggregate matrix together. Material distresses such as

fatigue cracking can be considered to be largely concentrated in the FAM. The shape and

gradation of coarse aggregates influence the overall performance of an asphalt mixture.

Since the focus of this study was to evaluate the influence of production temperatures on

the mechanical properties and performance of the mix, it is helpful to focus on the influence

of aggregate mineralogy by using FAM and filtering out the effect of shape and gradation

from coarse aggregates. The only requirement for this approach is that the fine aggregates

passing #16 sieve are selected to represent the mineralogy of the coarse aggregate in the

asphalt mixture.

The Dynamic Mechanical Analyzer (DMA) has been used in several studies to charac-

terize the aggregate-asphalt resistance to fatigue and moisture damage (Kim et al., 2004;

Branco et al., 2008; Masad et al., 2006). The DMA apparatus applies a cyclic, torsional

stress-controlled or strain-controlled load to a cylindrical sample of the FAM until failure

15

occurs. Specimens for DMA testing were prepared using fine aggregates (passing #16) and

asphalt binder in proportions that were representative of the original asphalt mixture.

Masad et al. (2008) used a comprehensive methodology to evaluate of the resistance

of DMA specimens to fracture under dry and wet conditions based on crack growth using

Paris law. They expressed Paris law in terms of the J integral of the dissipated pseudo-strain

energy as follows:

dr̄

dN= A [JR]

n (2.2)

where, r is the average crack radius in the specimen, and JR is the J-integral, which is the

pseudo-strain energy release rate per unit crack area defined below.

JR =∂WR

∂N

∂ (c.s.a)∂N

(2.3)

where, WR is the dissipated pseudo-strain energy (DPSE) per unit volume of the intact

material or the volume of the material that is capable of dissipating energy, c.s.a is the

crack surface area, which is equal to 2πr2 for a circular crack with radius r. By integrating

equation 2.2 the authors expressed the crack size as a function of loading cycles as follows:

r (N) =

(

2n+1n+1

) n+12n+1

(

A

(4πM)n

) 12n+1

N̂

N=0

(

∂WR

∂N

) nn+1

dN

n+12n+1

(2.4)

where, M is the number of cracks in a specimen. Based on the work of Schapery and Lyt-

ton, n was approximated as 1/m for strain-controlled testing where m is the exponent of

time in the power law equation for the relaxation modulus. Lytton et al. (2001) developed

a form for A as a function of viscoelastic properties, bond energy, and tensile strength.

These properties were computed based on surface energy measurements and DMA tests.

Masad et al. (2006) and Little and Bhasin (2006) demonstrated a good correlation between

moisture sensitivity of various asphalt mixtures observed in the field and the results ob-

tained using this approach with the DMA. Little and Bhasin (2006) also conducted tests on

dry and moisture conditioned specimen using the DMA in order to evaluate the moisture

sensitivity of the FAM.

Mixture performance can be characterized using a similar approach with the modifi-

cation that the tests will be conducted on asphalt mixtures instead of the FAM. Another

methodology to characterize mixture performance in terms of plastic deformation is based

16

on applying creep and recovery tests at relatively high temperatures. The results from

this test when analyzed using a viscoelastic-viscoplastic model provide information on the

recoverable and irrecoverable components of the mixture deformation and the damage ac-

cumulated in each of these components (Dessouky, 2005; Saadeh, 2005).

2.6 SUMMARY

In summary, this chapter presented a review of the various aggregate properties and their

significance in terms of dictating the strength, stability and moisture resistance of asphalt

mixtures. The chapter also presented a brief discussion on the various aspects of asphalt

mixture production that can influence these aggregate properties as well a few methods to

determine the aggregate properties and mixture performance. The scope of this study was to

(i) evaluate the influence of mixture production temperature on the adhesive characteristics

of aggregates and (ii) evaluate the adhesive characteristics of common minerals found in

aggregates. These two objectives are addressed in the following two chapters of this report.

17

CHAPTER 3. INFLUENCE OF MIXTURE PRODUCTION

TEMPERATURE ON ADHESIVE PROPERTIES OF

AGGREGATES

3.1 BACKGROUND

The production of hot mix asphalt (HMA) typically requires heating the asphalt binder

and aggregate to temperatures in the range of 140°C to 160°C. These high production

temperatures are required to generally achieve the required workability with the asphalt

binder and also to ensure that the aggregate surface is dry and forms a durable bond with

the asphalt binder. Therefore, selection of the mixing and compaction temperatures is

crucial to ensure proper aggregate coating, matrix stability during production and transport,

ease of placement, acceptable compaction, and ultimately acceptable performance of the

pavement.

Several new techniques have been developed to reduce the temperatures used for mix-

ing and compaction of hot mix asphalt (HMA). The asphalt mix produced at lower than

conventional mixing and compaction temperature using any of these techniques is referred

to as warm mix asphalt (WMA). Development of the WMA technology was initiated by the

German Bitumen Forum in 1997 in response to the Kyoto agreement. WMA is produced

at temperatures that are up to 40°C lower than the typical HMA production temperatures.

Lower production temperatures reduce emissions, fumes and odors, and hardening of the

binder during construction. Lower production temperatures also offer benefits such as en-

ergy savings, ability to open sites early and pave during cooler periods.

Some of the techniques currently being explored to produce WMA are: (i) creating a

foaming action within the asphalt binder at the time of mixing (e.g. WAM-Foam®), (ii)

use of emulsion to reduce binder viscosity (e.g. Evotherm®), (iii) use of mineral additives

to release water during the mixing process so that binder viscosity is reduced (e.g. Aspha-

Min® and Advera®), (iv) use of organic additives to reduce binder viscosity (e.g. Sasobit®

and Asphaltan B®), and (v) use of a combination of dry and wet aggregates during mix-

ing to release water and reduce binder viscosity (low energy asphalt). Irrespective of the

technique used to produce a WMA, a major concern while producing asphalt mixtures at

reduced mixing and compaction temperatures is inadequate drying of the aggregate surface

and its impact on the long term durability of the asphalt mixture.

19

Hurley and Prowell (2005a; 2005b; 2006) report that in some cases lower compaction

temperatures used to produce WMA increased the potential for moisture damage. They

speculated that incomplete drying of the aggregate at lower temperatures could have re-

sulted in the reduced durability of the asphalt mixture. Several researchers have evaluated

the performance of WMA produced using different techniques under laboratory conditions

(Hurley and Prowell, 2005a,b, 2006; Barthel et al., 2004; Sousa Filho et al., 2006; Wasi-

uddin et al., 2007). Most of these studies compare the performance of WMA to a similar

HMA. The findings from these studies enumerate potential durability and performance re-

lated issues that are associated with the use of different WMA technologies. However, it is

also noted that such issues have not been reported from recent field sections using WMA

mixes. This may be either due to the fact that such test sections are relatively few and recent

or that the laboratory tests do not accurately reflect the degradation processes in the field or

both.

There is a need to individually address each attribute that makes the WMA different

from a HMA. These include: i) efficiency of the technique or additive that is used to im-

prove binder and mixture workability, ii) impact of residual moisture within or on the sur-

face of aggregate on the performance and durability of the mix, and iii) long term impact

of the additive that remains as a part of the mix (after paving and compaction) on its per-

formance and durability.

The main objective of this study was to isolate and evaluate the impact of reduced mix-

ing temperatures on the residual moisture retained on aggregate surfaces. Residual moisture

retained on the aggregate surface can reduce the bond strength between the asphalt binder

and the aggregate as well as the wettability of the asphalt binder to the aggregate. Wetta-

bility and the bond strength between the binder and the aggregate are strongly correlated

to the performance and durability of the composite (FAM or asphalt mixture). A micro

calorimeter was used to measure the total energy of adhesion between the asphalt binder

and aggregates pretreated at different temperatures. The test method was devised such that

the measured total energy of adhesion was only affected by the pretreatment temperatures

of the aggregate.

In addition to the results from the micro calorimeter, mechanical tests were conducted

on Fine Aggregate Matrix (FAM) specimens produced at different mixing and compaction

temperatures. As discussed in section 2.5, the performance and durability of an asphalt

mixture is strongly related to the performance and durability of its FAM (Kim et al., 2004;

Arambula et al., 2007). FAM specimens were produced at three different mixing and com-

20

paction temperatures for different asphalt binder-aggregate combinations. The mechanical

behavior of the FAM was evaluated using a Dynamic Mechanical Analyzer (DMA). A ze-

olite based additive was used to produce FAM specimens at reduced temperatures. The

choice of zeolite rather than other additives was based on the following rationale. One

group of specimens was prepared at reduced mixing and compaction temperatures using

hydrated zeolite particles as filler. Addition of hydrated zeolite particles causes micro

foaming which in turn improves workability of the binder and allows for the production

of FAM specimens at reduced temperatures. However, during mixing and compaction the

zeolite particles dehydrate and are retained in the specimen as a filler. A second group of

specimens was prepared at conventional mixing and production temperatures. In this case

dehydrated zeolite particles were added as a filler to the mix. Use of dehydrated zeolite will

not cause foaming and consequently will not affect the workability of the mix at conven-

tional mixing and production temperatures. Therefore this approach allows one to produce

and compare two groups of mixtures that are identical in composition but are produced at

conventional and reduced mixing and production temperatures. Such comparisons are not

possible with the wax based or emulsion based additives.

3.2 MATERIALS

Three aggregates with different mineralogies were selected from the Strategic Highway

Research Program (SHRP) Materials Reference Library (MRL): Basalt (RK), Gravel (RL),

and Limestone (RD) (Robl et al., 1991). Two asphalt binders, AAB and AAD, were also

selected from the SHRP MRL. Binder AAB is a PG 58-22 grade from Wyoming and AAD

is a PG 58-28 grade California Coastal product (Jones, 1993). The selection of softer grade

asphalt binders renders the mechanical performance of the mixture more sensitive to the

aggregate properties at test temperatures (around 25°C).

3.3 INFLUENCE OF PRODUCTION TEMPERATURE ON ADHESION AND SUR-

FACE PROPERTIES OF AGGREGATES

3.3.1 Overview of the test method

A differential isothermal micro calorimeter (OmniCal Inc., USA) was used to measure the

influence of aggregate treatment temperature (used to represent production temperature)

on the adhesive properties of the aggregate. This was done using two different methods

of measurement. First the total energy of adhesion (TEA) between an asphalt binder solu-

21

tion and aggregate specimens treated at different temperatures were measured directly. The

TEA represents the total energy given out as heat when the liquid (binder) adheres to the

surface of the solid (aggreagte). Second, the surface free energy components of the aggre-

gate specimens treated at different temperatures were measured by using heat of immersion

(HoI) with three different probe liquids. The HoI is similar to the TEA, except that the total

heat is measured for aggregate-liquid pairs where in the liquid is a pure probe liquid with

known surface tension characteristics. The surface energy values calculated using the HoI

were then used to estimate the adhesive bond strength of aggregates treated at different

temperatures with typical asphalt binders.

The total energy of adhesion or heat of immersion is influenced by the specific surface

area (SSA) of the aggregates (the higher the SSA, the higher the energy released), and the

surface energy of the materials involved in the adhesion process. The SSA of the aggregates

in question was determined using a nitrogen adsorption method with BET equations (Gregg

and Sing, 1967).

The following is a brief description of the procedure to measure the TEA or HOI using

the micro-calorimeter. The solid or aggregate sample is placed in a 16 ml glass vial in

the reaction cell of the micro-calorimeter. Details of the solid sample are presented in

the following subsection. The glass vial has an open top polypropylene (PP) cap sealed

with a PTFE (poly-tetraflouroethylene) line silicone septa. A similar empty glass vial is

placed in the reference cell of the micro-calorimeter. Syringes with 21 gauge needles filled

with a probe liquid are placed over the silicone septa of the reaction and reference vials.

A temperature controlled aluminum jacket encloses the vials and the syringes within the

micro-calorimeter. A series of thermocouples interconnecting the reaction and reference

cell measures the differential heat flow between the two cells within an accuracy of 10

μwatts. After the cells reach thermal equilibrium, syringes with the probe liquid (or asphalt

solution) are punctured through the silicone septa and the liquid is injected in both cells

simultaneously. Immersion of the aggregate in the liquid causes heat flow between the

reaction cell and the reference cell, which is recorded by the data acquisition computer and

accompanying software. Thermal equilibrium between the reaction and the reference cell is

restored after the reaction is fully complete. Integration of the heat flow curve between the

two equilibrium states (before and after injection) yields the enthalpy of immersion. Note

that the term reaction does not necessarily imply a chemical reaction where two materials

react to form a third material. In this report, the term reaction is also used to denote the

physio-chemical surface interactions between the clean aggregate surface and the liquid.

22

Figure 3.1 illustrates a schematic of the device used in this study (Bhasin, 2006). A more

detailed description of the materials used and test procedure followed for each of the three

different tests is described in the following sections.

Figure 3.1. Schematic illustrating the working principle of the differential

micro-calorimeter (Bhasin 2006) with a reaction cell containing fine aggregates and

an empty reference cell; the syringes on cells are filled with the liquid used to

immerse the aggregate sample.

3.3.2 Sample preparation

Aggregates passing ASTM sieve #100 and retained on ASTM sieve #200 were used for

this test. The aggregates were washed with distilled water to remove dust particles from

surfaces, and dried overnight in an oven. Two vials were used for each test, one empty

(reference) and the other with 8 g ± 0.01 g of the aggregate (sample). Both vials were

subjected to the same conditioning procedure. The aggregates were treated at four differ-

ent temperatures (90ºC, 110ºC, 130ºC, and 150ºC). The heat treatment was carried out by

placing the sample and the reference vials on top of an electronically controlled heating

surface. The vials were heated with open caps for a period of three hours at the speci-

fied temperature. Immediately after removal from the heater, the vials were closed using a

poly-propylene cap with air tight silicone septa. The vials were then allowed to cool down

to room temperature. This procedure allowed the aggregate sample in the vial to retain

its surface characteristics, which were representative of different mixing temperatures. In

23

addition to the four treatment temperatures, samples of each aggregate type were prepared

at 150°C under vacuum for three hours as well as at room temperature (referred to as no

heat). High temperature and vacuum were used together in an attempt to create an ideal dry

aggregate surface. Additional details on the sample preparation procedure can be found in

(Vasconcelos et al., 2008).

In order to measure the TEA, stock solutions for different asphalt binders were prepared

using 1.5 g of asphalt dissolved in 11 mL of HPLC grade toluene. Previous studies con-

ducted at the Western Research Institute indicate that asphalt binder in a toluene solution

does not compromise the physio-chemical characteristics of the bitumen. Also, the bitu-

men molecules in a toluene solution have similar kinetics to those of molecules in liquid

bitumen at elevated temperatures (Western Research Institute, 2001). Figure 3.2 illustrates

the micro-calroimeter and a typical asphalt binder in the syringe that is injected into the

vial containing an aggregate sample (shown outside of the microcalorimeter).

The HoI of aggregates was measured using three different probe liquids: chloroform,

benzene, and heptane. All three liquids were HPLC grade. The HoI was used to estimate

the surface free energy of the aggregates.

3.3.3 Test and analysis

The micro calorimeter was used with its proprietary data acquisition and analysis software

to record heat flow during the test and to compute the total heat of immersion. The vial

with the aggregate sample was placed in the reaction cell and the empty vial was placed

in the reference cell of the micro calorimeter. Four syringes of 2 ml capacity each were

used to draw the asphalt solution or probe liquid. Two of these syringes were positioned

on top of the reaction vial and two on top of the reference vial (Figure 3.2). Heat flow

between the sample and reference cell was recorded using the software that accompanied

the micro calorimeter. The cells were allowed to reach thermal equilibrium. Equilibrium

was confirmed as the point when the heat flow ceased to change over time. After reaching

thermal equilibrium, the asphalt binder solution was injected into the vials.

The asphalt binder molecules preferentially adhere to the aggregate surface reducing

the total energy of the system and producing heat. The heat flow from the reaction cell was

recorded over time and the system was allowed to return to thermal equilibrium. The area

under the heat flow curve over time was integrated to obtain the measured TEA or HoI,

∆Hmeas (Figure 3.3). The magnitude of the TEA is proportional to the bond strength be-

tween the binder and the aggregate surface. In the case of the probe liquid, the magnitude

24

Figure 3.2. The micro-calorimeter (left) is shown with a typical vial containing an

aggregate sample (right); the asphalt binder is in solution form in the syringe ready

to be injected into the vial with the aggregate to measure the heat of immersion.

Figure 3.3. Typical heat flow measured using the micro-calorimeter when a solid

(aggregate) is immersed in a probe liquid.

25

of the HoI is proportional to the work of adhesion between the aggregate and the probe

liquid. A correction in the integrated value is necessary due to the difference in free vol-

umes inside the reaction and reference cells, ∆Hδv. The corrected enthalpy of immersion

is determined as follows:

∆Himm = ∆Hmeas−∆Hδv = ∆Hmeas−∆Hvapvs p0

RT(3.1)

where, vs is the volume occupied by aggregates in the vacuum sealed reaction cell, p0 is the

saturation vapor pressure of toluene (or probe liquid in the case of HoI) at the test temper-

ature, R is the universal gas constant, T is the test temperature, and ∆Hvap is the change in

enthalpy due to vaporization or heat of vaporization per mole of toluene or probe liquids.

A minimum of three replicates were tested for each combination of three aggregates and

six treatment conditions with the two asphalt binders and probe liquids.

For the interaction between asphalt binders and aggregates, ∆Himm divided by the total

surface area, Asur f ace, of the aggregates is the TEA or total energy of immersion per unit

surface area of the aggregate. The total surface area of the aggregates was determined as

the product of the specific surface area and the mass of the aggregate sample in the cell.

The specific surface area was measured using the multi-point BET method with nitrogen

as adsorbate, and with an outgas temperature of 150°C. The specific surface areas for RK,

RL, and RD were: 11.292; 3.803; and 0.906 m2/g, respectively. It must be noted that the

rank order as well as the order of magnitude for the specific surface area of therse three

aggregates is the same as reported in previous studies on these aggregates (Robl et al.,

1991).

For interaction between the three probe liquids and the aggregates, the following method

was used to estimate the surface free energy of the aggregate. Equation 3.2 relates the heat

of immersion, ∆Himm, to the three surface energy components of the aggregate and probe

liquid, and the total surface area of the aggregate, Asur f ace.

∆Himm = 2Asur f ace

(

γTotall −2

√

γLWl γLW

a −2√

γ+l γ−a −2√

γ−l γ+a

)

(3.2)

In equation 3.2, the superscripts LW , +, and − denote the non-polar, acid and base

components of surface free energy, respectively and the subscripts l and a denote the probe

liquid and aggregate, respectively. Since the specific surface areas of the aggregate were

measured using the nitrogen adsorption BET method, and the surface energy components

of the three liquids are known, the only unknowns in equation 3.2 are the three surface free

26

energy components of the aggregate. Consequently, heats of immersion with the three dif-

ferent probe liquids were used to generate a system of three linear equations (using equation

3.2) in order to determine the three surface free energy components of the aggregate.

3.3.4 Results

The two asphalt binders (AAB and AAD) were combined with the three aggregates (RK,

RL and RD) treated using six different conditions: at ambient conditions with no heat

treatment, heated and dried at 90°C, 110°C, 130°C, and 150°C in air, and heated and dried

at 150°C under vacuum. For the aggregates that were heated and dried in air, The TEA

was measured using three or more replicates for each of the aforementioned 36 combina-

tions. In addition, three probe liquids were used to determine the surface free energy of

aggregates as a function of the treatment temperature. Figure 3.4 illustrates the influence

of aggregate treatment temperature on the total energy of adhesion for the combinations of

different asphalt binders and aggregates. Figures 3.5, 3.6, and 3.7 illustrate the influence

of aggregate treatment temperature on the heat of immersion of these aggregates in differ-

ent probe liquids. The surface free energies of the aggregates were calculated using the

measured HoI with different probe liquids with equation 3.2. These surface energy values

were then used with the surface energy values for two different asphalt binders in order

to estimate the influence of aggregate treatment temperature on its adhesive bond strength

with the asphalt binder (Figure 3.8).

3.3.5 Discussion

One of the objectives of this study was to isolate the impact of reduced mixing temper-

atures on the residual moisture on aggregate surfaces and concomitant changes in the

binder-aggregate bond strength. This was achieved in two ways: (i) by determining the

TEA between asphalt binders and aggregates pretreated at different temperatures and (ii)

by measuring the surface energy components of pre-treated aggregates and using these to

compute the adhesive bond energy with asphalt binder. Results based on the direct mea-

surement of TEA and the computed energy of adhesion were similar and the following

discussion pertains to both.

Results indicate that the TEA between the asphalt binder and the aggregate either de-

creases slightly or is not significantly affected by the treatment temperature of the aggregate

within the temperature range of 90°C to 150°C (which is of interest for the production of

27

Figure 3.4. Influence of aggregate treatment temperature on the total energy of

immersion with asphalt binders.

Figure 3.5. Influence of aggregate treatment temperature on the heat of immersion

for aggregate RK.

WMA). The TEA between the asphalt binder and aggregate was also determined for aggre-

gates that were not subjected to any heat treatment. In this case, the TEA was significantly

28

Figure 3.6. Influence of aggregate treatment temperature on the heat of immersion

for aggregate RL.

Figure 3.7. Influence of aggregate treatment temperature on the heat of immersion

for aggregate RD.

lower compared to aggregates treated between 90°C to 150°C. This was as expected, since

aggregates that are not heated retain substantial amounts of surface adsorbed moisture and

result in reduced adhesion with the asphalt binder.

29

Figure 3.8. Influence of aggregate treatment temperature on computed adhesive

bond strength with different asphalt binders.

The conclusion would then, in summary, be that the TEA is substantially different for

aggregates that are not heated before mixing with asphalt binder and aggregates that are

heated before mixing with asphalt binder. However, the difference in TEA (for aggregates

mixed with asphalt binder) when the aggregates are heated to between 90°C and 150°C

before mixing with asphalt binder is not significant. A significant increase in the TEA

was observed for the limestone aggregate when the aggregate were treated at 150°C under

vacuum as compared to when it was treated at the same temperature without vacuum. This

increase was not substantial for the other predominantly siliceous aggregates.

The results from this study were compared to the findings from other similar studies

on different types of silicate minerals. Ligner et al. (1989) presented the variation in the

dispersive components of the surface energies for different silicates (amorphous and crys-

talline) as a function of thermal treatment. A decrease in the dispersive component of the

aggregate surface energy will reduce its bond strength with any given asphalt binder. Con-

sequently this should result in a lower TEA when measured using the micro calorimeter.

The results reported by Ligner et al. (1989) for crystalline and amorphous silicates suggest

that the TEA (at least for the siliceous aggregates) should not change significantly for the

temperature range used in this study (90 to 150°C). This was consistent with the results

30

based on the TEA measured using the micro calorimeter.

3.4 INFLUENCE OF PRODUCTION TEMPERATURES ON PERFORMANCE OF

FAM

3.4.1 Sample preparation

The mixture design for the Fine Aggregate Matrix (FAM) followed the method described

in Zollinger (2005). Fine aggregate particles smaller than 1.18 mm in size (passing ASTM

sieve #16) were used in the mixture design. The fine aggregates were separated into differ-

ent size fractions and recombined to match the gradation of a typical dense graded mixture.

The fine aggregates were oven dried for four hours at the selected mixing temperature prior

to use. The aggregates were then mixed with the asphalt binder and aged for a period of

two hours at the selected compaction temperature. Each combination of asphalt binder and

aggregate was prepared at three different mixing and compaction temperatures.

The first mix was prepared using a mixing and compaction temperature representative

of conventional hot mix asphalt (according to ASTM D 2493). The other two mixtures

were prepared at 20°C and 40°C below conventional mixing and compaction temperatures.

These two mixtures are representative of WMA. Zeolite was added at the rate of 0.3% by

weight of the total mixture during the preparation of each of the three mixtures. As de-

scribed previously, the most important difference was that for the conventional hot mix,

the zeolite was dehydrated by heating it to the same temperature as the aggregate prior to

mixing. Since dehydrated zeolite does not release water during mixing it can be treated as

a mineral filler that is incorporated within the FAM. For the mixes prepared at reduced tem-

peratures, hydrated zeolite stored at room temperature was added to the aggregates along

with the asphalt binder during mixing. In this case, hydrated zeolite particles come into

contact with the binder, at relatively higher temperature. This causes the zeolite particles

to release their moisture and the binder to foam simultaneously improving the workability

of the mixture.

After mixing and short term aging, the dehydrated zeolite particles can be considered

to remain in the mix as mineral filler. The Superpave Gyratory Compactor (SGC) was

used to compact the mix and fabricate samples approximately 80 mm height in a 150 mm

diameter mold. The compaction was set to stop at 87% of the maximum specific gravity

(Gmm). The ends of the 150 mm diameter sample were sawed to achieve a target height

of 50 mm. Following this, approximately 15 specimens each of 12 mm diameter were

31

obtained by coring the 150 mm diameter SGC specimen. The bulk specific gravity of the

DMA specimens was determined and the average air void content of the samples was 15%

(+/- 1.5%). The specimens were then glued using epoxy to metal holders for testing with

the DMA.

3.4.2 Test and analysis

All tests using the DMA were conducted in the controlled-strain mode. A torsion load was

applied following a sinusoidal wave form at a constant frequency (10 Hz) and temperature

(24°C) for all the tests. The test was performed in two steps: (i) a low strain amplitude

(0.0065%) was applied to obtain the linear viscoelastic complex modulus of the material;

and (ii) a high strain amplitude (0.15%) was applied to induce damage in the specimens.

The progression of damage during the cyclic loading process was monitored to determine

the fatigue damage resistance of the FAM. Both steps were conducted on at least four

replicate specimens for each of the eighteen mixtures, and at least four more replicates were

tested following moisture conditioning. Moisture conditioning was carried out by partially

saturating the specimens (65 - 80%) using a vacuum pump, and leaving the specimens

under water at room temperature for a period of 12 hours.

At high strain amplitudes, the apparent or measured shear modulus of the specimen

gradually decreases as the number of load cycles increases until the specimen fails. The

results from the DMA were analyzed to obtain three different measures of performance

for the FAM: (i) the undamaged viscoelastic properties of the FAM quantified based on its

shear modulus (G*) measured at 1200 cycles at a low strain amplitude, and (ii) the fatigue

damage resistance of the FAM determined based on a dissipated strain energy parameter.

Dissipated energy was used as a measure of fatigue life as it typically has much lower

variability compared to the number of load cycles to failure (Branco, 2008; Branco et al.,

2008). This was also confirmed based on the data from this study. In addition, the mois-

ture damage resistance of the FAM was quantified by comparing the above parameters for

moisture conditioned specimens to the same parameters for unconditioned specimens. The

following dissipated strain energy parameter was used to characterize the fatigue cracking

resistance of the FAM specimens.

W =12

γ20(

G∗V E −G∗N f

)

(3.3)

where, G∗V E is the complex shear modulus of the FAM specimen in its initial state deter-

32

mined using the low strain amplitude (0.0065%), G∗N f is the shear modulus just before

specimen failure when subjected to the cyclic test at the high strain amplitude (0.15%),

and γ0 is the target constant strain amplitude during the fatigue test (0.15% in this case).

The number of load cycles to failure were identified from a plot of NG∗1/G∗N where N is

the number of load cycles, G∗1 is the complex shear modulus at the beginning of the test

and G∗N is the complex shear modulus at the Nth cycles. The number of load cycles at the

maximum value of NG∗1/G∗N was considered as N f (Figure 3.9). In addition to W from

equation 3.3, other dissipated energy parameters may also be used to characterize fatigue

damage in the FAM specimens. Masad et al. (2007) present the use of other parameters

such as the crack growth index to characterize the fatigue cracking life of FAM specimens.

In the context of this study, the material properties required to compute the crack growth

index could not be readily determined for specimens produced at the different tempera-

tures. Furthermore, since the objective of this study was to evaluate the impact of mixing

and compaction temperature on the mechanical behavior of FAM specimens, the parameter

based on equation 3.3 was considered to be adequate to characterize the fatigue cracking

resistance for comparative purposes.

Figure 3.9. A typical plot used to identify the number of load cycles to failure; note

that curve with the number of load cycles multiplied by the normalized reduction in

shear modulus has a well defined peak that is identified with failure.

To determine the moisture damage resistance of the mixtures, the specimens were par-

33

tially vacuum saturated (65-80%) and left under water for a period of 12 hours at room

temperature. After removal from water, the specimens were tested again the same strain

amplitude, temperature, and frequency as the unconditioned specimens. The shear modu-

lus at low strain amplitude and the dissipated energy parameter, W , at high strain amplitude

were determined as before. For each material combination, the values of the shear modulus,

G∗, and dissipated energy parameter, W , were normalized with respect to results from the

unconditioned mix prepared at 155°C. This facilitated the comparison of both factors (pro-

duction temperature and moisture conditioning) on a common scale. In other words, for

each material combination the normalized values of G∗ or W indicate the relative change

in the parameter due to the reduction in mixing temperature and/or moisture conditioning

with respect to the unconditioned mix prepared at 155°C.

3.4.3 Results

In the first step of the DMA test, shear modulus of the FAM specimens was determined

by applying 1200 load cycles (2 minutes at 10Hz) in torsion by applying a low strain am-

plitude. Figure 3.10 illustrates the average and standard deviation of the complex shear

modulus of mixtures prepared at different mixing temperatures; results were obtained on

three or more replicates for each material. In the second step of the DMA test, fatigue

cracking resistance of the FAM specimens was determined by applying cyclic loading at

a higher strain amplitude until failure occurred. The dissipated energy parameter W was

computed for all replicates of the 18 mixtures using equation 3.3. Figure 3.11 illustrates the

results for the average and the ± one standard deviation of the dissipated energy parameter,

W . A higher value of this parameter indicates better fatigue cracking resistance. Figures

3.12 and 3.13 illustrate the impact of mixing temperature and moisture conditioning on G∗

and W , normalized for each of the six different material combinations.

3.4.4 Discussion

The mechanical properties of the unconditioned FAM specimens, as determined using the

DMA, either decreased with a decrease in the mixing temperature or did not change signifi-

cantly (Figures 3.10 and 3.11). This is most likely due to the differences in the effectiveness

of the warm mix additive to enable mixing and compaction at reduced temperatures. An-

other, although less likely, factor that can contribute to these differences is the increase in

residual moisture on the aggregate surface resulting in poor bonding with the asphalt. The

34

Figure 3.10. Influence of mixing and compaction temperature on the shear modulus

of dry specimens (average shown with +/- one standard deviation).

Figure 3.11. Influence of mixing and compaction temperature on the fracture

resistance of dry specimens (average shown with +/- one standard deviation).

influence of this factor was isolated and discussed in section 3.3.5. Finally, the interaction

of the binder with the WMA additive, in this case the zeolite particles, is the least likely

factor for the differences in the mechanical properties shown in Figures 3.10 through 3.13.

35

Figure 3.12. Influence of moisture conditioning on linear viscoelastic property of

different FAM specimens.

Figure 3.13. Influence of moisture conditioning on fracture resistance of different

FAM specimens.

This is because the zeolite particles were added to all mixtures, WMA as well as those

produced at conventional mixing and compaction temperatures. It must be noted that the

aforementioned statement does not imply that additive does not interact with the binder. In

other words, for the experiment design used in this study, such interactions were normalized

36

by using a dehydrated form of the additive in the control mix.

Another aspect that must be noted from these results is that, for most combinations

of materials and mixing temperatures, moisture conditioning resulted in a reduced value

of G∗ and W as compared to the unconditioned specimens. In most cases, the response

of the moisture conditioned specimens prepared at different mixing temperatures followed

the same trend as that of the unconditioned specimens. Another observation was that, for

any given mixing temperature, the FAM prepared using asphalt binder AAD showed better

or similar resistance to moisture induced damage compared to asphalt binder AAB. This

is contradictory to findings from previous studies that indicate that asphalt binder AAD

has poor resistance to moisture induced damage (Little and Bhasin, 2006). The anomalous

moisture resistance of materials with the binder AAD may be explained as follows. Ad-

dition of divalent cations to the asphalt binder has been shown to result in the formation

of water insoluble salts with carboxylic acids within the mastic. This promotes the migra-

tion of other functional groups such as pyridines, ketones, and sulfoxides to the aggregate

surface in order to form more durable, moisture resistant adhesive bonds. The above mech-

anism has been used to explain the improved moisture damage resistance of certain asphalt

binders with the addition of hydrated lime (Little and Petersen, 2005). Asphalt binders

with more polar and asphaltene content are generally most responsive to the action of the

hydrated lime (Plancher et al., 1976; Petersen, 1984). Since the asphalt binder AAD has

high asphaltene content and polar functional groups (Jones, 1993), and the zeolite that was

used in this study was rich in calcium ions, the authors hypothesize that the high moisture

damage resistance of binder AAD may be due to the mechanisms described above.

37

CHAPTER 4. PHYSIO-CHEMICAL AND ADHESIVE

PROPERTIES OF COMMON MINERALS IN AGGREGATES

4.1 BACKGROUND

As discussed previously, surface energy of bitumen and aggregate can be effectively used

to calculate bond strengths and energy parameters to rank the susceptibility of mixtures

to incur damage due to moisture. Prior to this research a comprehensive study of surface

energy characteristics within and across mineral groups had not been undertaken nor such

data reported. Therefore, the primary purpose and objective of this part of the study was to

assemble and report such data. Although aggregates used in asphalt mixtures are seldom

comprised of a single mineral, understanding the bond strength between bitumen and ag-

gregates or mineral surfaces (dry and in the presence of moisture) is essential to understand

the variability and source of adhesive bond strength between the aggregates and the binder.

Such a database can lead to a better understanding of minerals to avoid in asphalt mixture

combinations.

4.2 FACTORS THAT INFLUENCE SURFACE FREE ENERGY OF AGGREGATES

4.2.1 Non-polar active sites

Chemical surface energy components are located on natural surfaces in sites, commonly

termed ‘active sites’. These are the positions where sorption occurs. Combining known

knowledge of individual mineral surface characteristics to surface energy results can cor-

relate type and density of active surface sorption sites on minerals. The sorption sites on

minerals are of varying type (Johnston, 2002). There are two major positions that an active

site of a mineral can be found. These are on the edges and the basal surfaces. Of these two

positions an active site can be either polar or non-polar. Non-polar sites primarily bond with

van der Waals components of adsorbates. Non-polar sites are commonly found on micas,

zeolites, kaolinite, serpentine minerals, smectites, and a variety of other silicate minerals.

Vermiculites and chlorites have non-polar sites to a lesser degree (Johnston, 2002). Micas

and zeolites often have the strongest non-polar sites. non-polar sites are generally found on

the neutral siloxane surface of silicates. These surfaces have no charge and no permanent

dipole moment. Therefore, they are termed hydrophobic. Because of this hydrophobic na-

39

ture water has little or no interaction with neutral siloxane surfaces and they are not able to

form hydrogen bonds. Hydrophobic surfaces occur on 2:1 phyllosilicates where no isomor-

phic substitution has occurred. This is a result of the -2 charge on the oxygen atoms being

completely satisfied by neighboring silicon atoms (Huheey et al., 1983). Siloxane surfaces

act as very weak Lewis bases but are mostly inert. Their surface energy is therefore domi-

nated by van der Waals forces. For this reason, although water and other polar molecules

have a slight affinity for neutral siloxane surfaces, non-polar organic solutes and non-polar

regions of larger biological molecules such as proteins and enzymes can efficiently bond to

this type of surface through van der Waals interactions (Johnston, 2002).

4.2.2 Polar active sites

Polar active sites on surfaces can be broken down into permanently charged sites, condi-

tionally charged sites, exchangeable metal cations, and exposed uncoordinated metal atoms

(Johnston, 1996). Polar sites react with Lewis acid/base components of adsorbates. Polar

sites are generally termed hydrophilic as a result of their charge and dipole moments. Con-

ditionally charged sites are pH dependent. Some examples of pH dependent sites are on

iron oxides such as goethite, aluminum oxides such as gibbsite, manganese oxides such as

birnessite, palygorskite and sepiolite, and a large number of other silicate minerals (John-

ston, 2002). In addition to these kaolinite, serpentine, phyrophyllite, talc, micas, zeolites

such as analcime, carbonates, and even titanium and zirconium minerals have some condi-

tionally charged surfaces.

Conditionally charged sites are primarily a result of inorganic surface hydroxyl groups.

They must always be located on edge sites. The best example of this is gibbsite as explained

by Johnston (2002). Permanent charge sites are often found on aluminosilicates such as

allophane (contains pH dependent sites as well), micas such as biotite, silicate clays such as

montmorillonite (also contain pH dependent sites, and vermiculite, and zeolites. Kaolinite

and serpentinites also contain some permanently charged sites.

Constant charged sites are a result of isomorphic substitution and defects within the

mineral. Isomorphic substitution is a process where one atom or molecule is replaced

by another of similar size. Isomorphous substitution often occurs in either the octahedral

or tetrahedral sheets of 2:1 phyllosilicates. This switch is sometimes accompanied by a

change in charge by the substituted atom. Constant charge sites are characterized by a

permanent negative charge (Johnston, 1996). One example is the substitution of Mg2+

for Al3+ in the octahedral sheet of layer silicates. Al3+ often substitutes for Si4+ in the

40

tetrahedral sheet. Both of these result in a gain of one electron. Isomorphic substitution

can result in a localized or delocalized charge. Mg2+ for Al3+ is thought to be delocalized

over approximately nine oxygen atoms on the siloxane surface whereas Al3+ for Si4+ is

more concentrated at around three oxygens (Sposito, 1989).

Closely related to constant charged sites described above are the exchangeable metal

cation sites. In this case, however, the metal cation is not replaced by another cation. In-

stead, the organic solute coordinates directly to the cation occupying the isomorphic substi-

tution site (Isacsson and Sawhney, 1983; Johnston, 1996). One example of this occurrence

is when phenols react with exchangeable alkali and alkaline earth metal cations on these

sites. The degree of metal-organic solute attraction will be a function of the organic solutes

ability to compete for coordination sites around the metal center (Isacsson and Sawhney,

1983). It is also possible for exchangeable and structural transition metals in their oxi-

dized state to interact directly with organic solutes. These metal cations can act as Lewis

acids when they accept electrons from the organic solute (Johnston, 1996; Voudrias and

Reinhard, 1986). An example of this reaction is when Cu2+ or Al3+ interact with reduced

aqueous solutes. Cu2+ and Al3+ are themselves reduced and a radical organic cation is

produced on the surface. These last two types of polar active sites are common on clay

minerals.

4.2.3 Coatings

All surface energy measurements are based on the surface conditions at the time of mea-

surement(Parks, 1990). Any surface can be altered to change its surface energy (Staszczuk,

1985). An example of this is when large organic molecules sorb to phyllosilicates. The

organic molecule balances the charges of the polar and non-polar sites. The interface es-

sentially becomes part of the interior of a new material comprised of the mineral and the

organic molecule (Neu, 1996). The new edge site has very different characteristics than the

layered silicate had originally. Smaller changes can also affect the surface energy charac-

teristics. If a metal cation is sorbed from an aqueous solution onto the surface of a mineral

it can satisfy a net negative charge imbalance. This reaction causes the negative component

to decrease. Conversely a loss of a metal cation leads to an increase in the electron donor

component of the surface free energy. In addition the presence of water or other aqueous

solution at the surface will increase the positive component.

Water also transports various media making sorbate species available to the material

surface. Conditionally charged surfaces are also strongly affected by the external environ-

41

ment. Acidic solution provides excess hydrogen positive charges to the surface of the min-

eral increasing the electron acceptor component. Basic solutions remove positive charges

from the mineral surface causing the electron donor component to increase. Thus, the ex-

ternal environment can alter the surface energy even when coatings do not bond to the

surface.

4.3 MATERIALS AND METHODS

The method to establish the pure phase minerals’ surface energies involves using the Uni-

versal sorption device (USD). This device carries the advantage of being a convenient

method that can be used on a routine basis with minimal human biases. All tests were

performed on pure mineral samples that were purchased or obtained. The surfaces of the

pure phase minerals were characterized using an electron microprobe. This characteriza-

tion involves establishing the surface mineralogy and purity of each mineral. Tables 4.1

and 4.2 list the pure minerals used in this study.

4.3.1 Elemental analysis of the minerals

Elemental analyses were performed on the bulk properties of the minerals. Measurements

were achieved by crushing the samples below a number 35 standard sieve and placing them

in an oven for 24 hours at 75°C. The samples were then placed in a desiccator and per-

cent nitrogen, sulfur, and carbon was measured with an Elemental Analyzer. Sulphanilic

acid was used to establish the calibration curve at .072, .56, 1.04, 1.6, and 2.6 mg. Sul-

phanilic acid composition was measured at 8.09% nitrogen, 41.61% carbon, and 18.5%

sulfur. Montana soil standards reference material (SRM) 2711 was used for quality control

every 25 samples.

Blanks were placed after each Montana soil standard to allow venting of residual ma-

terial from the standard and the previous samples. All samples were analyzed in quadru-

plicate unless otherwise noted. Detection limits were set at 0.01 N%, 0.018% C, and 0.01

S% based on the calibration curve used for these analyses, however personal experience

on this device has shown that these detection limits are conservative. The minerals were

washed with distilled (DO) and reverse osmosis (RO) water prior to crushing to remove

large particles on the surface. Aggregates were also washed with DO and RO water prior

to crushing. After crushing, the aggregates were separated and three batches were made

for all except RD-7 Limestone. Batch number one was placed in consecutive hydrochloric

42

acid solutions until effervescence ceased. Batch one was performed in quadruplicate and

placed in the oven. Batch two was placed in consecutive hydrochloric acid solutions in

quadruplicate and not placed in the oven. Batch three was not placed in acid solutions or in

the oven. This was done to quantify organic carbon and carbonate carbon in the samples.

No nitrogen was measured above detection limits in any of the samples. Both carbon and

sulfur were present above detection limits in most samples. Carbon to nitrogen ratios are

listed for comparison. The detailed results from the elemental analysis of the minerals used

in this study can be found (Miller, 2009).

4.3.2 Pretreatment of minerals for surface energy testing

Samples of the minerals were crushed with a stainless steel impact mortar to be passed

through a #4 (4.74 mm) ASTM standard sieve and retained by a #8 (2.36 mm) ASTM

standard sieve. This method was chosen as opposed to saws and pressed powders to mimic

natural erosion by breaking along cleavage or fracturing. After the samples were crushed

they were be washed with reverse osmosis water to remove large particles and then with

distilled water. The minerals were then placed in an oven (Fisher Isotemp 200 series) at

75-80°C for at least one day. The samples were then allowed to cool for one hour and then

were carried to the USD. They were not placed in a desiccator to cool so that they would

continue to closely model minerals in the environment.

4.3.3 Testing with the universal sorption device

Theoretical background

As discussed in the previous sections, the USD indirectly measures the surface free energy

of solid surfaces based on the work of adhesion of a solvent vapor onto the solid. The

USD uses adsorption isotherms to calculate the specific surface area of the solid, and the

spreading pressure of the vapor over the solid. Hexane vapor is used to calculate the specific

surface area of the mineral using the relationship:

A =

(

nmN0

M

)

α (4.1)

where A is the specific surface area, N0 is Avogadro’s number, M is the molecular weight

of the probe vapor (86.16 for hexane), α is the projected area of a single molecule, and

nm is the monolayer capacity of the solid surface (Gregg and Sing, 1967). Bhasin and

43

Little (2006) used the liquid density formula to estimate the projected area of a hexane

molecular as 39Å2. However, Bhasin found that using 56.2Å2, which was a value found in

the literature, fit Micrometrics N2 specific surface area measurements of aggregates more

tightly (McClellan and Harnsberger, 1967). The estimated 56.2Å2 hexane molecule will

hereafter be referred to as the standard hexane molecule.

The monolayer capacity is the number of molecules required to cover the solid surface

in a single layer and is calculated as:

nm =1

I +S(4.2)

where S is the slope and I is the intercept of the best fit line between P/n(P0−P) and P/P0 ;

and where P is the partial vapor pressure, P0 is the maximum saturation vapor pressure, and

n is equal to the mass of vapor adsorbed per unit mass of the mineral. From the isotherm

the equilibrium spreading pressure of the reference vapor over the solid is calculated using:

πe =RT

MA

pnˆ

0

n

Pd p (4.3)

where πe is the equilibrium spreading pressure, R is the universal gas constant, T is the test

temperature, M is the molecular weight of the probe vapor, n is the mass of vapor adsorbed

per unit mass of the mineral at the vapor pressure P, and A is the specific surface area of the

mineral. The spreading pressure is related to the work of adhesion through the equation:

WSN = πe +2γTotalv (4.4)

where WSN is the work of adhesion between the reference vapor and the solid, πe is mean

equilibrium spreading pressure of the reference vapor over the solid, and γTotalv is the total

surface energy of the reference vapor. Finally, using the Good - van Oss theory the work of

adhesion can be related to surface energy using the equation:

WSN = 2√

γLWs γLW

v +2√

γ+s γ−v +2√

γ−s γ+v (4.5)

where WSN is the work of adhesion between the reference vapor and the solid being tested

and other terms are as described before where subscript v refers to the probe vapor and

subscript s refers to the solid (aggregate or mineral). Using three probe vapors with known

characteristics (n-hexane, methyl propyl ketone, and water), a system of equations can

44

be created in order to solve for the three unknowns in equation 4.5. For a more detailed

description of the USD device and its usage the reader is referred to Bhasin and Little (6).

Testing

After the samples were oven dried approximately 20 to 25 grams of a mineral sample was

placed in the sample cage, and the sample column was closed. The sample and sample

column were then brought to vacuum (< 0.05 mbars) and hot degassed at 60°C for at least

two hours. Next, the Rubotherm magnetic suspension balance was run continuously (auto

balance) until the readings remained stable within 0.001g. When the analytical balance

remained consistently stable n-hexane was introduced into the sample chamber. Hexane

has no electron acceptor or electron donor components; therefore it only exerts van der

Waals forces onto the mineral sample. For this reason, hexane can be used to calculate the

van der Waals component for the mineral surface energy. The sample was then exposed to

ten equal increments of partial vapor pressure from vacuum to saturation vapor pressure.

After each increment the adsorbed mass was recorded to plot the adsorption isotherm after

it reached equilibrium. The process was repeated using methyl propyl ketone (MPK) and

water as the reference vapors. For each mineral, the adsorption isotherms using the three

reference vapors were used to compute the surface free energy components of the mineral

sample as well its specific surface area following the procedure described above.

4.4 RESULTS FROM SURFACE ENERGY TESTING

4.4.1 Carbonates

Five carbonates were analyzed on the USD: Calcite (CaCO3), Cerussite (PbCO3), Siderite

(FeCO3), Rhodochrosite (MnCO3), and Dolomite (CaMg(CO3)2). The carbonate series

was most useful for comparing surface hardness and softness. This was accomplished by

comparing the surface chemistry of the various cations. The SSA of the carbonates ranged

from .06 to .28 cm2/g based on the standard hexane molecule. Methylpropylketone was

very low and did not vary much across the carbonate series.

Calcite and Dolomite

Calcite surface energy was primarily non-polar with a fractional polarity of .25. The total

surface energy was 46.54 ergs/cm2. Dolomite surface energy was also non-polar with a

fractional polarity of .25. Dolomite and calcite differed considerably on the van der Waals

component, however with 34.94 and 60.29 ergs/cm2 respectively. The polar surface energy

45

of calcite was 11.6 ergs/cm2 while dolomite was approximately double at 20.28 ergs/cm2.

Rhodochrosite, Cerussite, and Siderite

The polar surface energy increased from cerussite, rhodochrosite, to siderite. This increase

corresponded with the hard/soft acid base concept. The van der Waals energy was 35.07,

40.33, and 61.39 ergs/cm2 for cerussite, rhodochrosite, and siderite. Cerussite was limited

in sample quantity. For this reason an error analysis could not be performed. The coefficient

of variation for the total surface energy of rhodochrosite was 5.81% and 16.21% for siderite.

Siderite was the only sample to have a fractional polarity greater than .50.

4.4.2 Sulfates

Gypsum was acquired for analysis. In order to improve any heterogeneity the bassanite and

hot gypsum were prepared from the gypsum sample. After the gypsum was fractured with

the impact mortar between a number 4 and 8 sieve 1/3 of the sample was placed in the oven

at 75°C for two weeks. When gypsum (CaSO4•2H2O) is heated to 70°C it loses 1.5 waters

and becomes Bassanite (commonly known as hemihydrates CaSO4•0.5H2O).

The rest of the sample was placed in a plastic bag and stored. Half of the remaining

sample was placed in vacuum and hot degassed in the same manner at the other minerals

while the rest was allowed to get to vacuum but not hot degassed. The bassanite measured

the highest SSA with all the reference vapor calculations. Hot gypsum measured higher

SSA forMPK and both waters indicating higher affinity for polar molecules. The difference

was most profound with water. Bassanite SSA for 5 and 10 square angstrom water was

approximately 30 times higher than hot gypsum and 40 times higher than gypsum. Gypsum

measured higher SSA for both hexane molecular sizes.

Bassanite

The bassanite sample showed a large increase in polar surface energy (59.90 ergs/cm2). The

van der Waals component was slightly lower than gypsum and hot gypsum. Fewer sample

runs were performed on bassanite than the other two sulfates based on time constraints on

the universal sorption device. This caused an increase in standard error. The coefficient of

variation for the standard hexane molecule was 29.19% for the total surface energy.

Gypsum and Hot Gypsum

As discussed in the chemistry section the gypsum sample was homogeneous and the sample

size was over 500 grams. The coefficient of variation for the standard hexane molecule was

46

1.81% for non-polar and 8.9% for polar surface energy. The 95% confidence interval for

the total surface energy was 56.45 to 62.84 ergs/cm2. Hot gypsum measured similar results

to gypsum. The main differences were increased affinity to polar probe vapors. The polar

surface energy of hot gypsum was 21.49 while the polar surface energy for gypsum was

18.52 ergs/cm2. The fractional polarity was 3% higher.

4.4.3 Phyllosilicates

Four phyllosilicates were analyzed on the sorption device: biotite, muscovite, kaolinite,

and montmorillonite. Two of the minerals were clays (kaolinite, and montmorillonite).

These samples were separated from all the other minerals based on their specific surface

areas. SSAmeasurements with the standard hexane molecule were two orders of magnitude

higher than the other mineral specimens. Kaolinite SSA was 10.53, and montmorillonite

SSA was 23.30. These agreed well with literature values.

Muscovite & Biotite

Muscovite and biotite are 2:1 phyllosilicates with ¼ of the tetrahedral sites occupied by

Al3+ instead of Si4+. This charge increase is balanced by monovalent potassium in the

interlayer. The main difference between biotite and muscovite is that biotite is trioctahedral

and muscovite is dioctahedral. This means that the hydroxides in the octahedral sheet are

balanced by a divalent cation in trioctahedral sheets and a trivalent cation in dioctahedral

sheets. This is the reason that biotite is more dense than muscovite (3.09 and 2.82).

The total surface energy for muscovite and biotite were 82.07 and 67.41 ergs/cm2.

The non-polar component was roughly equal with biotite begin slightly higher (47.55 and

52.51), but muscovite had a larger polar component (34.52 to 14.90). The acid component

of biotite was the second lowest of the mineral set (.01). This increased the error of this

measurement due to its proximity to the detection limit. The coefficient of variation for the

acid component of biotite was 181.49%.

Kaolinite

It was also necessary to place the kaolinite on the special holder inside the sorption device

because of the friable nature of kaolinite. The total surface energy of kaolinite was 70.51

ergs/cm2. The magnitude was mainly due to the polar component, 40.02 ergs/cm2. This

gave a fractional polarity of .57. Thus, the polar component of kaolinite was similar to

montmorillonite, but the non-polar component was much less. The acid component was

47

higher than the acid component for montmorillonite, but both were higher than the mineral

average.

Montmorillonite

The montmorillonite mineral sample had a non-polar surface energy of 42.85 ergs/cm2.

This was higher than kaolinite but lower than the two micas. The polar component was

22.45 ergs/cm2. This was lower than kaolinite and muscovite. The montmorillonite was

homogeneous, but very friable. For this reason the sample had to be placed in a specially

designed holder cage inside the USD in order to catch the sample if any broke off during

adsorption of the probe vapors. The error for the sample reflected the homogeneity. The

coefficient of variation of the total surface energy was 10.66% and the 95% confidence

interval was 51.65 to 78.94 ergs/cm2.

4.4.4 Feldspars

Four feldspar samples were sampled: microcline, albite, andesine, and labradorite. Micro-

cline is a Potassium rich Feldspar. Albite, andesine, and labradorite are all members of the

plagioclase Na-Ca series. Albite has the most sodium (90-100%). Andesine and labradorite

are intermediate composition (50-70 & 30-50% respectively). Feldspars are very common

igneous, sedimentary, and metamorphic minerals, and therefore are very important to the

study. The specific surface areas of the feldspars ranged from .10 to .27 cm2/g. This range

was higher than the average of the other mineral groups and very similar to the aggregate

specific surface areas. Labradorite had the highest SSA for every reference vapor except

water.

Albite

The albite sample was from a pegmatite and homogeneous. This was displayed in the

tight non-polar surface energy range. The coefficient of variation for the standard hexane

molecule was 3.54%. The 95% confidence interval was 48-55.15 ergs/cm2 with an average

of 51.57. The acid component was the lowest of the feldspars (.22), but the base component

was the second highest second to andesine).

Andesine

The andesine sample measured the highest total surface energy of the feldspars (129.88

ergs/cm2). This was mainly due to the polar component measured at 89.24 ergs/cm2. The

fractional polarity was also the highest of the feldspars at .69.

48

Labradorite

The chemical composition of labradorite and andesine were very similar (Na.45Ca.50Al1.51Si2.50O8

and Na.49Ca.45Al1.47Si2.53O8 respectively both with minor potassium). Thus, any differ-

ence in surface energy might be attributed to other characteristics such as surface coatings

or surface morphology. The total surface energy was measured at 82.92 ergs/cm2 with

non-polar and polar surface energies of 46.21 and 36.71. Thus, the fractional polarity was

.44. The acid component was much higher than the average for the feldspars (1.81 and .72

ergs/cm2). The base component, however was the lowest of the feldspars.

Microcline

The microcline (KAlSi3O8) was perthitic. Perthitic feldspars are inter grown with sodic

alkali feldspar. Perthitic texture and composition is very common in feldspars. The sample

was pegmatitic which increased the value of the mineral as a homogeneous sample for

measuring the surface energy. Microcline had a large van der Waals component of 44

ergs/cm2, and the polar component was only 19.35. This gave a fractional polarity of .31.

The acid component was .46 ergs/cm2. This was very close to andesine and albite. The

total surface energy was 63.35 ergs/cm2.

4.4.5 Oxides

The oxides measured were quartz (SiO2), hematite (Fe2O3), and ilmenite (FeTiO3). Quartz

is also a tectosilicate similar to the feldspars sampled. Specific surface areas for the oxides

ranged from .05 to .1 cm2/g for the standard hexane molecule. Using water as the SSA

vapor lowered the measurement slightly. For a 10 square angstrom water molecule the

SSA ranged from .03 to .04 cm2/g. Using MPK as the SSA probe vapor lowered the

measurements again to a range of .01 to .02 cm2/g.

Hematite

The total surface energy of hematite was measured as 128.81 ergs/cm2. This was higher

than the other oxides. The sample had higher polar surface energy than non-polar with

a fractional polarity of .62. The sample was homogeneous but limited in amount. This

caused the confidence interval to be larger than desired. There was enough sample to give

an estimate of the magnitudes of the components.

49

Ilmenite

The ilmenite sample was also limited in quantity; however the homogeneity of the sample

(see chemistry section) caused the surface energy measurements to remain tight in range.

The total surface energy 95% confidence interval was 49.5-72.2 ergs/cm2 with a coefficient

of variation of 9.47%. The average total surface energy for the standard hexane molecule

was 60.89. The non-polar surface energy was 39.76 was 21.13 ergs/cm2 giving a fractional

polarity of .35. This showed that the sample surface energy was controlled by the non-polar

portion.

Quartz

The quartz van der Waals component measured 50.33 ergs/cm2 for the standard hexane

molecule and 68 ergs/cm2 for 39Å2 hexane molecule. The acid component for quartz

was close to zero. This measurement was close to the detection limit for the sorption

device. This caused the coefficient of variation for the acid component and the total polar

component to be much larger than other minerals. The coefficient of variation for the

acid component was 471%, and the coefficient for the total polar component was 72.81%.

The quartz sample surface energy was controlled by the non-polar component (fractional

polarity of .09). Thus, there were not many unfilled bonds on the surface of the sample.

4.4.6 Nesosilicates and Inosilicates

The neso/Inosilicate group had the largest variation in surface energies of any mineral

group. Augite had the highest surface energy at 367.78 ergs/cm2. The fractional polar-

ity was .95. The specific surface areas of this group were modest, however with a range

of .03 to .25 cm2/g. The van der Waals component increase from olivine, hornblende, to

augite. The fractional polarity of olivine was .30. Heterogeneity increased from olivine,

hornblende, to augite.

A summary of the surface energy results is presented in Table 4.3 and Figure 4.1.

4.5 DISCUSSION

In the context of asphalt mixtures, the surface free energies of mineral aggregates are im-

portant to quantify the moisture damage or stripping potential of different aggregate-binder

combinations. The correlation between the surface properties of these materials and their

tendency to strip in the presence of water is relatively well established in the literature.

50

Figure 4.1. Summary of total, polar and non-polar components of surface energies

for minerals (arranged by group and in the order discussed above).

The three quantities based on the surface energies of asphalt binders and aggregate that are

related to the moisture sensitivity of an asphalt mixture are:

• work of adhesion between the asphalt binder and aggregate (WAB),

• work of debonding or reduction in free energy of the system when water displaces

asphalt binder from a binder-aggregate interface (WABW ), and

• work of cohesion of the asphalt binder or mastic (WBB).

The above three quantities are computed using the surface free energy components of the

individual materials. For an asphalt mixture to be durable and have a relatively low sensi-

tivity to moisture, it is desirable that the work of adhesion, WAB, between the asphalt binder

and the aggregate be as high as possible. Furthermore, the greater the magnitude of work

of debonding when water displaces the asphalt binder from the binder-aggregate interface,

WABW , the greater the thermodynamic potential that drives moisture damage will be. There-

fore, it is normally desirable that this quantity be as small as possible. This is because this

term is usually (almost always) a negative quantity, and the more negative it is, the more

likely it is for water to replace bitumen at the interface with mineral or aggregate. Little

and Bhasin (2006) developed an energy ratio (ER) based on the two energy terms WAB

and WABW and which combines these terms into a single, dimensionless parameter. They

51

later adjusted this ratio to also consider the affinity of the bitumen to develop a cohesive

interaction (WBB). The resulting energy ratio (ER2) is expressed by:

ER2 =WAB−WBB

WABW

(4.6)

Equation 4.6 defines a parameter that can be used to estimate the moisture sensitivity of

asphalt mixtures based on the hypothesis that moisture sensitivity is directly proportional

to the dry adhesive bond strength, and inversely proportional to the work of debonding or

the reduction in free energy during debonding. Another form of this equation used in the

earlier study was the product of ER2 and specific surface area of the aggregate. The work

of debonding is determined as follows:

WABW = γAW + γBW − γAB (4.7)

where, γAW is the bond strength between the aggregate and water, γBW is the bond strength

between bitumen and water, and γAB is the interfacial bond energy between the bitumen and

the aggregate. Figures 4.2 and 4.3 show evidence of the utility of equation 4.6 in predicting

moisture damage based on laboratory experiments (tensile modulus, Figure 4.2, and plas-

tic deformation, Figure 4.3) with twelve different mixtures (twelve combinations of three

aggregates and four bitumens). In the tensile test, Figure 4.2, the ratio of tensile resilient

modulus in the wet (moisture conditioned) state compared to the tensile resilient modulus

in the dry state is plotted on the ordinate versus ER2 on plotted on the abscissa. In Figure

4.3, repeated compressive loading was continued until the sample accumulated one percent

permanent strain. The ratio of the number of cycles to accumulate one percent strain for

each mixture under wet (moisture conditioned) and dry conditions is plotted on the ordinate

versus the energy ratio (equation 4.6) on the abscissa. In both figures, the dashed lines rep-

resent the 95 percent confidence range and demonstrates the strong relationship between

moisture damage, as recorded by these ratios, and ER2.

Consider the energy ratio of the form shown in 4.6 for the different minerals with two

substantially different asphalt binders (AAB and AAD). The properties of these binders

were obtained from previous studies (Hefer et al., 2006). The energy ratio was computed

for all combinations of these two binders tested with the minerals. Figure 4.4 shows these

energy ratios for each mineral evaluated as the ERs progress from low to high (most resis-

tant to moisture damage). The ratios are calculated for the two asphalts (AAB and AAD)

with the ER always higher (more moisture resistant) for bitumen AAB, regardless of the

52

Figure 4.2. Laboratory performance (tensile modulus) of mixtures compared to ER2

* SSA (ER2 is a parameter based on surface free energies of asphalt binder and

aggregate and SSA is the specific surface area of the aggregate). (After Little and

Bhasin, 2006).

mineral considered. Since surface area, diffusivity (sorption of bitumen into pores spaces,

etc.) also affect bond strength, the ERs do not necessarily mean superior resistance to mois-

ture damage. However, they should be an excellent indicator of the inherent tenacity of the

bond based on adhesive bond strength.

The efficacy of the data in Figure 4.4 is supported by Figure 4.5, which compares the

wet to dry fatigue ratio for bitumens AAB and AAD for three different aggregates. In a

series of controlled experiments on dense graded asphalt mixtures, it was demonstrated that

a combination of AAD with three different types of aggregates was relatively more suscep-

tible to moisture induced damage as compared to AAB. These results represent the ratio

of fatigue life of moisture conditioned specimens to dry specimens. The fatigue tests were

conducted in direct tension mode by applying cyclic loads with constant stress amplitude

at a frequency of 10Hz. All test specimens were designed and fabricated to have a similar

binder content, air void content, and gradation.

Finally, Figure 4.6 presents the energy ratio data by mineral category and shows by the

number in the text block the number of minerals within the group for which the calculated

53

Figure 4.3. Laboratory performance (plastic deformation) of mixtures compared to

ER2 * SSA parameter (ER2 is a parameter based on surface free energies of asphalt

binder and aggregate and SSA is the specific surface area of the aggregate) (After

Little and Bhasin, 2006).

value of was positive. A positive value of indicates that the bitumen developed a preferred

(thermodynamically stable) adhesive bond with the mineral surface compared to water (one

phyllosilicate, two carbonates, one sulfate and one neo-silicate). More specifically, the min-

erals that formed inherently thermodynamically stable bonds with bitumen were olivine

in the neo / ino silicate group, gypsum and hot gypsum in the sulfate group, calcite and

cerussite in the carbonate group, and montmorillonite in the phyllosilicate group. In other

words, these mineral surfaces form a bond with bitumen that is more stable than the bond

with water suggesting that stripping due to water is not thermodynamically favorable for

these combinations. Hundreds of measurements of aggregate and bitumen surface energies

at Texas A&M University indicate that none of the aggregate-bitumen combinations ana-

lyzed thus far were thermodynamically resistant to stripping. This apparent contradiction is

resolved by considering the fact that aggregates used to produce asphalt mixtures comprise

of more than one mineral. It is very likely that the sum effect of all minerals on the ag-

gregate surface will usually result in an aggregate-bitumen bond that is thermodynamically

prone to stripping as evidenced in the previous studies.

54

Figure 4.4. Energy ratio representing inherent resistance to moisture induced

damage for mineral-asphalt binder pairs presented from lowest to highest and for

asphalts AAB and AAD (higher energy ratio indicates better resistance to moisture

damage).

Surface energy results from this study on pure minerals indicate that there is a possi-

bility that localized regions on aggregate surfaces with these minerals may be inherently

resistant to stripping. However, such minerals are very limited in number and the stability

would depend on the mineral as well as the surface chemistry of the binder. It is also inter-

esting to note that no minerals within the oxides and feldspar group were found that were

thermodynamically stable with either one of these two binders. Certainly, these results

deserve more evaluation and consideration.

Although the prime objective of this research was to assess the variability of bond

strength developed between asphalt and the 22 representative minerals selected, a sec-

ondary objective was to understand the complex interrelationship between the three master

variables of surface energy: surface chemistry, surface morphology, and surface coatings.

However, the complex interactions among these master variables and the indirect manner

in which surface energies of solids must be measured prevents a clear path to separating

out the clear impact of these master variables. Nonetheless, several trends were observed.

55

Figure 4.5. Comparison of fatigue lives of moisture conditioned asphalt mixtures to

dry asphalt mixtures for bitumen AAB and AAD.

Figure 4.6. Surface energy trends within a mineral group and identification of

minerals that preferred a bond with bitumen rather than water.

First, organic coatings appear to have the greatest impact aside from the mineral type,

and they have the most impact when the amount becomes great enough to be considered

56

as a separate phase. This is most evident in the comparison of carbonates. The hard/soft

acid/base (HSAB) concept appears to hold true until the amount of organic carbon on the

surface increases significantly. The siderite sample (4.4 moles of organic carbon) appears

to have enough organic coating to have established a separate phase. None of the other

carbonate samples had as much organic carbon (1.07, 1.95, 0.92, and 1.46 moles). How-

ever, a weak positive trend can also be seen when all of the Lewis Base component surface

energies are plotted against moles of surface organics (Figure 4.7).

Second, organic coatings may increase the Lewis base component. By comparing each

of the individual mineral groups there is no noticeable pattern that arises from the magni-

tude of the organic material on the surface with van der Waals and Lewis Acid. However,

when comparing all of the groups together there appears to an increase in Lewis base sur-

face energy with organic coatings. A weak linear correlation exists among each of the

mineral groups between Lewis basicity and total organic carbon. This might show that

surface coatings always play a role in surface energy, but are additionally affected by other

variables. If a strong correlation was seen then it could be argued that organic coatings are

a dominant control on Lewis Basicity. This, however was not seen.

Third, the universal sorption device specific surface area calculations correlate well

with literature values for similar minerals. Bhasin and Little (Bhasin and Little, 2006)

found that a good correlation exists between aggregate SSA measurements of n-Hexane

adsorption and BET nitrogen adsorption. This research found that results compared well

with published values for similar minerals. This may be evidence that the assumption of

n-Hexane preferentially laying along its long axis is correct.

57

Figure 4.7. Base surface energy versus moles of organic carbon on mineral surfaces.

58

Table 4.1. List of Minerals for Surface Energy Measurements

Minerals Group Formula Comments Acquisition

Andesine Tectosilicate Na(70-

50%)

Ca(30-

50%)

(Al,Si)

AlSi2O8

Dominant feldspar in andesite. Minor

in granite and metamorphic rocks.

TAMU

Collection

Albite Tectosilicate NaAlSi3O8 Found in granite and metamorphic

rocks.

Maine

Augite Inosilicate (Ca,Na)(Mg,Fe,Al)

(Si,Al)2O6

An important rock-forming mineral in

many igneous rocks, especially in

gabbros and basalts.

TAMU

Collection

Bassanite Sulfate CaSO4•

0.5H2O

An active mineral that constitutes

Plaster of Paris.

RNG

Collection

Biotite Phyllosilicate K(Mg,Fe)

3(AlSi3O10)

(OH)2

Biotite is a common rock forming

mineral present in most igneous rocks

and both regional and contact.

TAMU

Collection

Calcite Carbonate CaCO3 Common in sedimentary rocks.

Calcite is also the predominant

mineral in limestone, the predominant

rock that is crushed into aggregate.

Mexico

Cerussite Carbonate PbCO3 An ore of lead. Tsumeb,

Namibia

Dolomite Carbonate Ca

Mg(CO3)2

A common sedimentary rock-forming

mineral, dolomitic limestone.

RNG

Collection

Gypsum Sulfate CaSO4•

2H2O

Common mineral found in arid

landscapes

RNG

Collection

Hematite Iron Oxide Fe2O3 Formed as a secondary weathering

mineral in soils.

RNG

Collection

Ilmenite Iron

Titanium

Oxide

FeTiO3 Common oxide in igneous

environments

RNG

Collection

Kaolinite Clay Al2Si2 O5

(OH)4

Common clay found in variety of soils

and aggregates

RNG

Collection

Labrad-

orite

Tectosilicate Ca(50-

70%)

Na(50-

30%)

(Al,SI)

AlSi2O8

Labradorite is a common feldspar. Naim,

Labrador

59

Table 4.2. List of Minerals for Surface Energy Measurements

Minerals Group Formula Comments Acquisition

Microcline Tectosilicate KAlSi3O8 Common feldspar found in granites. TAMU

Collection

Montm-

orillonite

Clay (Na,Ca)

(Al, Mg)

6(Si4O10)

3(OH)6

Common clay found in variety of soils

and aggregates.

RNG

Collection

Muscovite Phyllosilicate KAl2(AlSi3O10)

(F,OH)2

Common silicate in igneous,

sedimentary, and metamorphic

environments.

RNG

Collection

Olivine Nesosilicate (Mg,Fe)

2SiO4

Found in ultramafic igneous rocks and

marbles that formed from

metamorphosed impure limestones.

San

Carlos,

Arizona

Quartz Tectosilicate SiO2 Most abundant mineral of the crust;

ubiquitous in all environments.

Arkansas

Rhodochrosite Carbonate MnCO3 Minor ore of manganese. RNG

Collection

Siderite Carbonate FeCO3 Common mineral found in

sedimentary formations.

Idaho

60

Table 4.3. Summary of Surface Energy Measurements of Minerals (ergs/cm2)

Mineral Acid Base Total

Polar

Compo-

nent

van der

Waals

Total

Surface

Energy

Andesine 0.40 3755.04 77.70 40.64 118.34

Albite 0.22 501.69 21.22 51.57 72.79

Augite 8.69 3890.33 367.78 52.67 420.45

Bassanite 0.30 3036.03 59.90 38.27 98.16

Biotite 0.07 809.97 14.90 52.51 67.41

Calcite 0.40 85.16 11.60 34.94 46.54

Cerussite 0.11 113.14 7.04 35.07 42.11

Dolomite 0.18 564.05 20.28 60.29 80.57

Gypsum 1.31 65.47 18.52 41.13 59.65

Gypsum (HD) 1.32 87.66 21.49 42.24 63.73

Hematite 2.85 558.07 79.82 48.99 128.81

Ilmenite 0.35 318.90 21.13 39.76 60.89

Kaolinite 5.01 80.00 40.02 30.48 70.51

Labradorite 1.81 186.54 36.71 46.21 82.92

Microcline 0.46 202.79 19.35 44.00 63.35

Montmorillonite 1.57 80.43 22.45 42.85 65.29

Muscovite 0.55 544.68 34.52 47.55 82.07

Olivine 1.55 57.52 18.87 44.17 63.04

Quartz 0.02 365.00 5.04 50.33 55.37

Rhodocrosite 0.86 145.76 22.33 40.33 62.66

Siderite 1.59 789.63 70.80 61.39 132.18

61

CHAPTER 5. CONCLUDING REMARKS

5.1 INFLUENCE OF MIXTURE PRODUCTION TEMPERATURE

The use of warm mix asphalt has several benefits such as energy savings and reduced

emissions. Several research studies have compared the performance of HMA to a similar

WMA produced using different additives and techniques. While these comparisons are

typically made on a holistic basis, there is a need to individually assess the source and

impact of the various factors that differentiate a WMA from a HMA. The objectives of this

study were to evaluate the impact of reducing mixing temperatures on the: i) mechanical

behavior of the fine aggregate matrix (FAM) portion of an asphalt mixture, and ii) residual

moisture on aggregate surfaces and concomitant adhesive bond strength with the asphalt

binder.

In this study, a calcium based synthetic zeolite was used to prepare FAMs produced

over a range of mixing and compaction temperatures. The key findings from this study are

as follows:

• Results based on the dynamic mechanical analysis of the FAM specimens indicate

that for a given material combination, shear modulus and fatigue cracking resis-

tance of the FAM typically decreased when mixing and compaction temperatures

decreased.

• Results based on the micro calorimeter demonstrate that for the aggregates with rel-

atively low porosity used in this study, lower aggregate pretreatment temperatures

(within the range of 90oC to 150oC) did not significantly impact the total energy of

adhesion. From this, the researchers infer that residual moisture on the aggregate

surfaces at reduced mixing temperatures (within the range of 90oC to 150oC) did not

significantly contribute to the observed reduction in the shear modulus and fatigue

cracking resistance at lower mixing temperatures.

• A link between the diminished mechanical properties of the FAMs based on DMA

measurements and the reduction in pretreatment temperatures of the aggregate and

the rate of dehydration of the zeolite filler was established. Moisture egress from

zeolite was determined to be slower at lower treatment temperatures. It is speculated

that this could have resulted in the potential for more residual moisture in mixtures

63

prepared at lower temperatures. The higher residual moisture content of the FAMs

prepared at lower temperatures is consistent with inferior mechanical properties.

• Based on the historical performance data of the binders used in this study, it appears

that the use of a synthetic zeolite with calcium ions may help improve the moisture

damage resistance of certain asphalt binders.

An important consideration related to the above findings is that the findings are limited to

aggregates with low porosity. For porous aggregates, such as certain types of limestone,

lowering the mixing temperature may influence the moisture absorbed within the aggregate

bulk and consequently contribute to the adverse performance of mixtures. The authors are

conducting further research on the impact of using porous aggregates with WMA.

5.2 PHYSIO-CHEMICAL AND ADHESIVE PROPERTIES OF MINERAL AGGRE-

GATES

The USD can differentiate among surface energy measurements of aggregate and mineral

surfaces, and such measurements can be used to calculate bond strengths between bitu-

men and aggregate and to identify mixtures susceptible to moisture damage. Twenty-two

minerals that commonly comprise the surfaces of aggregates were carefully selected and

tested. The authors believe this to be the most extensive catalog of mineral surface proper-

ties ever compiled. This catalog can serve as a reference for properties that can be used to

explain specific interactions between minerals and materials interacting with these miner-

als including but not limited to bitumen and water. Surface energies vary over the range of

minerals tested; and within a group or category of aggregate, substantial differences in sur-

face energy occur. These differences are probably related to chemical composition, surface

roughness or surface area, and coatings. Due to the complexity of the interactions of the

master variables, the authors could not, independently validate the impact of each master

variable. However, a trend between organic coating and surface energy was identified.

The dimensionless energy ratio of adhesive bond strength between aggregate (mineral)

and bitumen in dry condition to the bond strength in the presence of water demonstrated a

wide variation based on the mineral considered and always favored the bond with bitumen

AAB compared to bitumen AAD and this is verified by laboratory mixture testing. How-

ever, it was surprising that some minerals developed a thermodynamically favored bond

with bitumen compared to the competing bond between aggregate and water. Previous

testing on aggregates and computation of energy ratios always revealed the thermodynamic

64

potential of water to replace bitumen at the aggregate surface. However, aggregates are not

generally purely comprised of one mineral type, and the large majority of the minerals

tested demonstrated a potential to strip. This finding requires more evaluation.

65

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