a towards the revision of austroads design of pavements with cemented materials

AP-T167/10

AUSTROADS TECHNICAL REPORT

Towards the Revision of Austroads Procedures for the Design of Pavements

Containing Cemented Materials

Towards the Revision of Austroads Procedures for the Design of Pavements Containing Cemented Materials


Published September 2010

Austroads Ltd. 2010

This work is copyright. Apart from any use as permitted under the Copyright Act 1968, no part may be reproduced by any process without the prior written permission of Austroads.


ISBN 978-1-921709-39-5

Austroads Project No. TT1358

Austroads Publication No. APT167/10

Project Manager Allan Jones

Prepared by

Geoff Jameson

Published by Austroads Ltd. Level 9, Robell House 287 Elizabeth Street

Sydney NSW 2000 Australia Phone: +61 2 9264 7088

Fax: +61 2 9264 1657 Email: [email protected]

www.austroads.com.au

Austroads believes this publication to be correct at the time of printing and does not accept responsibility for any consequences arising from the use of information herein. Readers should

rely on their own skill and judgement to apply information to particular issues.


Sydney 2010

Austroads profile Austroads purpose is to contribute to improved Australian and New Zealand transport outcomes by:

providing expert advice to SCOT and ATC on road and road transport issues facilitating collaboration between road agencies promoting harmonisation, consistency and uniformity in road and related operations undertaking strategic research on behalf of road agencies and communicating outcomes promoting improved and consistent practice by road agencies. Austroads membership Austroads membership comprises the six state and two territory road transport and traffic authorities, the Commonwealth Department of Infrastructure, Transport, Regional Development and Local Government, the Australian Local Government Association, and New Zealand Transport Agency. Austroads is governed by a Board consisting of the chief executive officer (or an alternative senior executive officer) of each of its 11 member organisations:

Roads and Traffic Authority New South Wales Roads Corporation Victoria Department of Transport and Main Roads Queensland Main Roads Western Australia Department for Transport, Energy and Infrastructure South Australia Department of Infrastructure, Energy and Resources Tasmania Department of Lands and Planning Northern Territory Department of Territory and Municipal Services Australian Capital Territory Department of Infrastructure, Transport, Regional Development and Local Government Australian Local Government Association New Zealand Transport Agency. The success of Austroads is derived from the collaboration of member organisations and others in the road industry. It aims to be the Australasian leader in providing high quality information, advice and fostering research in the road sector.


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CONTENTS

1 INTRODUCTION ...................................................................................................................... 12 1987 AUSTROADS GUIDE ...................................................................................................... 23 VICROADS ADOPTED 8TH POWER RELATIONSHIP ............................................................ 43.1 Original Analysis and Conclusion ............................................................................................. 4

3.1.1 Laboratory Fatigue Relationship ................................................................................ 43.1.2 Field Fatigue Relationships ....................................................................................... 5

3.2 Further Analysis ...................................................................................................................... 113.2.1 Modulus Dependency .............................................................................................. 11

4 1997 CHANGE TO 12TH POWER RELATIONSHIP ............................................................... 175 RECENT RESEARCH FINDINGS .......................................................................................... 185.1 Austroads Project TT1065 Findings ....................................................................................... 18

5.1.1 Laboratory Testing ................................................................................................... 185.1.2 Fatigue Under Accelerated Loading ........................................................................ 21

5.2 Austroads TT1359 Findings .................................................................................................... 256 DISCUSSION ......................................................................................................................... 277 CONCLUSIONS ..................................................................................................................... 28REFERENCES ................................................................................................................................ 29


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TABLES

Table 3.1: Models to predict fatigue life to half modulus .............................................................. 8Table 3.2: Models to predict fatigue life to 1.0 m/m2 cracking ................................................... 11Table 5.1: Summary for flexural moduli data ............................................................................. 18Table 5.2: Summary for breaking strain data ............................................................................. 19Table 5.3: Summary of models characterising the mean reduction in modulus for

each axle load ........................................................................................................... 22Table 5.4: Summary of Load Damage Exponents (LDEs) based on mean data at

each axle load ........................................................................................................... 22Table 5.5: Modulus, breaking strain and fatigue relationships for each material ....................... 26

FIGURES

Figure 3.1: Laboratory fatigue testing of field beams .................................................................... 5Figure 3.2: Reduction in cemented materials modulus with accelerated loading ......................... 6Figure 3.3: Fatigue relationships to half initial cemented material modulus ................................. 7Figure 3.4: Predicted strain correlation with back-calculated modulus ......................................... 8Figure 3.5: Fatigue relationships to 1 m/m2 surface cracking ...................................................... 10Figure 3.6: Predicted dependence of fatigue on cemented material modulus ............................ 11Figure 3.7: Comparison of back-calculated moduli with measured field core and

beam values ............................................................................................................. 12Figure 3.8: Reanalysis of accelerated loading data .................................................................... 13Figure 3.9: Prediction error variation with cemented material modulus ...................................... 14Figure 3.10: Factors influencing the variation in predicted strain .................................................. 14Figure 3.11: Relationship for sites with initial modulus less than 10 000 MPa.............................. 15Figure 3.12: Relationship for sites with initial modulus greater than 10 000 MPa......................... 16Figure 5.1: Crushed hornfels laboratory fatigue results of laboratory beams ............................. 20Figure 5.2: Crushed siltstone laboratory fatigue results of laboratory beams ............................. 20Figure 5.3: Crushed hornfels fatigue results under accelerated loading: data lives

103 to 107 .................................................................................................................. 23Figure 5.4: Crushed siltstone fatigue results under accelerated loading: data lives

103 to 107 .................................................................................................................. 24Figure 5.5: Crushed hornfels fatigue results under accelerated loading: data with

initial moduli greater than 8000 MPa ........................................................................ 24Figure 5.6: Crushed siltstone fatigue results under accelerated loading: data with

initial moduli greater than 10 000 MPa ..................................................................... 25


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SUMMARY The laboratory characterisation of cemented materials for road pavements has recently been completed in Austroads Project TT1359 Cost-Effective Structural Treatments for Rural Highways. This project concluded there is a need for a substantial revision to the thickness design procedures for cemented materials as provided in Austroads Guide to Pavement Technology Part 2: Pavement Structural Design (Austroads 2010).

This report provides a research strategy for this revision. It reviews:

the origins of the current design procedures results of past and recent laboratory and accelerated loading trials on cemented materials

modulus and fatigue.

Based on this review a research strategy is proposed to deliver the revised Austroads procedures.


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1 INTRODUCTION The laboratory characterisation of cemented materials for road pavements has recently been completed in Austroads Project TT1359 Cost-Effective Structural Treatments for Rural Highways. This project concluded there is a need for a substantial revision to the thickness design procedures for cemented materials as provided in Austroads Guide to Pavement Technology Part 2: Pavement Structural Design (Austroads 2010).

This report provides a research strategy for this revision. In particular:

Section 2 outlines the origins of the 18th power fatigue relationship adopted in 1987 Austroads Guide (NAASRA 1987).

Section 3 describes the finding of the 1990/91 Mulgrave accelerated loading trial of a high quality Victoria cement treated crushed rock and the associated laboratory testing which lead to VicRoads adopting an 8th power fatigue relationship in 1993.

Section 4 outlines the origins of the 12th power fatigue relationship adopted in the Austroads Guide in 1997 and modified for project reliability in the 2004 Guide.

Section 5 summarises the findings of the two recent Austroads research projects. Section 6 summarises the issues highlighted by this research review, issues which need to

be addressed in amending the design procedures in Austroads Guide to Pavement Technology Part 2: Pavement Structural Design (Austroads 2010).


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2 1987 AUSTROADS GUIDE Part 1 Section 4.3.2 of Austroads Technical Report AP-T33-04 Technical Basis of Austroads Pavement Design Guide (Austroads 2008b) describes the origins of the fatigue relationship adopted in the Austroads Pavement Design Guide (NAASRA 1987).

At the time of developing the Guide, the leading proponent within Australia of extensive use of cemented materials was the Main Roads Department, Queensland (MRDQ). Apace with this increased use, an increased appreciation of its performance developed within MRDQ. The Working Group (WG) availed itself of this knowledge base and was guided by them in the formation of fatigue relationships for cemented materials.

The performance relationship in both the 1987 and 1992 versions of the Guide is (Equation 1):

N = (K/)18 1

where

N = is the number of repetitions of tensile strain at the bottom of the cemented layer before fatigue failure occurs, i.e. when the level of this strain is microstrain.

The numerator K depends on the stiffness of the material as follows:

Modulus of cemented material (MPa) Value of K

2 000 280

5 000 200

10 000 150

In the development of pavement thickness design curves for Queensland conditions based on elastic analysis methods, Baran and Aubrey (1978) noted the following relationship (Equation 2) developed by Pretorius (1969).

N = (142/)20.3 2

The modulus of the material was not known, but assumed to be > 10 000 MPa (later confirmed in Pretorius and Monismith (1972) to be 28 000 MPa). They also noted a (graphical) relationship between strain at break and modulus for cement-treated natural weathered gravel in Walker et al. (1977).

The Pretorius relationship gave a tolerable strain level of 72 for 106 repetitions which, from the Walker et al. plot corresponded to 65% of the strain at break for materials stiffer than 10 000 MPa. This ratio (tolerable strain for 106 repetitions)/(strain at break) = 0.65 was adopted as being applicable to materials with moduli down to 2000 MPa. For a given modulus, the corresponding strain at break was determined from the Walker et al. plot and then multiplied by 0.65 to give the tolerable strain for 106 repetitions. In a similar manner, values of this ratio for 105 and 107 repetitions were determined for the Pretorius material and applied to less stiff material.


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On this basis, fatigue relationships were developed for materials of moduli 2000, 5000, 7000, and 10 000 MPa over the range 105 to 107 strain repetitions. These relationships were then used in the development of thickness design charts by Baran and Aubrey (1978). Angell (1988) reported that the relationships for the materials with moduli 2000, 5000 and 10 000 MPa were of the form (Equation 3):

N = (K1/)K2 3with the values of K1 and K2 as follows:

Modulus of cemented material (MPa) K1 K2

2 000 259 19.9

5 000 244 14.5

10 000 152 18.3

Litwinowicz (1982) undertook a review of the basis for these relationships and found that (in Angells words):

the general level of these relationships appeared to be appropriate but that their exact form and slope still required further investigation.

The WG, in reviewing these relationships, expressed some surprise that the value of the exponent (K2) did not change monotonically with the material modulus.

Further investigations were undertaken and the relationships eventually adopted by the WG for inclusion in the 1987 Guide were recommended by Litwinowicz (1984, private communication) on the basis of his investigations.

Subsequent to the WGs adoption of the relationships with exponent 18, Angell (1988), in the course of development of a pavement design manual for MRDQ, undertook a further review of the literature and reported fatigue exponents of 32, 9, 12.7, 12.2 and 12 for relationships developed by workers in four countries.

In light of this, together with his proposition that, if the true exponent were 18, then cement-treated pavements in Queensland would be failing very early in their design life because of vehicle overloading, Angell opted for an exponent of 12 and derived numerator (K1) values such that the revised relationships (in his words) allow approximately the same levels of strain as the relationships previously used. Angell developed the following relationship (Equation 4):

N = (K/)12 4

with values of K as follows:

Modulus of cemented material (MPa) K1 K2

2 000 259 19.9

5 000 244 14.5

10 000 152 18.3

The WG was apprised of MRDQs intention to adopt these revised relationships while the original (1987 version) Guide was in press.

The WG decided it was too late to make the change to the 1987 Guide, and it was not until 1997 that the Guide was changed to a 12th power relationship as discussed in Section 4.


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3 VICROADS ADOPTED 8TH POWER RELATIONSHIP

3.1 Original Analysis and Conclusion In 1990/91 Austroads investigated the fatigue performance of cement treated crushed rock (Jameson et al. 1992). The project included both laboratory fatigue testing and accelerated loading of 3% cement treated crushed rock, along the alignment of a proposed arterial in Mulgrave, Melbourne. The accelerated loading was undertaken using the Accelerated Loading Facility (ALF).

3.1.1 Laboratory Fatigue Relationship Laboratory third-point repeated flexure testing of a field beam was undertaken after more than 12 months field curing in the test pavements. The fatigue life was defined as the number of cycles when the modulus had decreased to half the initial value, these being determined after 25 cycles of bedding-in. The mean initial flexural modulus of these field beams was about 5300 MPa.

The fatigue data are plotted in Figure 3.1. In reporting the fatigue results Jameson et al. (1992) mentioned that initially regression analysis was undertaken with the logarithm of the strain as the independent variable (x axis) and the logarithm of the fatigue life as the dependent or observed variable (y axis). This is the correct procedure to enable fatigue life to be estimated from strain provided the error in strain is substantially less than the error in fatigue life. This relationship is also shown in Figure 3.1.

As argued by Jameson et al. (1992) and more recently confirmed by a more rigorous orthogonal analysis of fatigue data (Gonzalez et al., 2010), undertaking the regression with the logarithm of the fatigue life as the independent variable and the logarithm of strain as the dependent variable is more appropriate. As shown Figure 3.1, the strain dependency of this latter relationship is about 8, considerably greater than when fatigue life is the dependent variable.

Consequently, the laboratory fatigue testing of the field beams indicated fatigue life was inversely related to about 7-8 power of applied strain. Note that the slope was similar when beams with fatigue lives less than 1000 were deleted.

The average breaking strain of the field beams was about 600 microstrain, whereas for laboratory compacted beams the average was about 145 microstrain; the latter is consistent with recent laboratory testing of laboratory compacted beams (Section 5.2). The much higher and variable breaking strain of field beams was also reported in another recent Austroads Project TT1065 (Section 5.1). This difference in the characteristics of laboratory beams and in-service cemented material is an important consideration in the future strategy to revise the Austroads design procedures.

These high breaking strains of the aged field beams suggest that the fatigue results may have been influenced by micro-cracking. This may explain why fatigue results are less dependent on strain (strain exponent about 8) than observed in recent test results of laboratory compacted and cured beams, (average strain exponent 22 as described in Section 5.2).


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140

160

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200

220

240

1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06

Load repetitions to half initial modulus

Tens

ile s

trai

n (m

icro

stra

in)

dataLife dependent variableStrain dependent variable

Source: Jameson et al. 1992.

Figure 3.1: Laboratory fatigue testing of field beams

3.1.2 Field Fatigue Relationships Over 2 million cycles of accelerated loading were applied to 14 field experiments using dual wheel loading of 40 kN, 60 kN and 80 kN. To develop fatigue relationships, for each chainage of each experiment the initial layer moduli were back-calculated from the measured Falling Weight Deflectometer deflections. In terms of the cement treated crushed rock, the mean initial back-calculated modulus was 10 900 MPa, significantly higher than the mean flexural moduli of the field beams (5300 MPa): this modulus difference may have been due to the variability of modulus along the test section (Section 3.2.1, Figure 3.7). Using these layer moduli, the initial strains under the applied loading were predicted.

To develop fatigue relationships, these initial strains need to be related to the observed cemented materials fatigue life. This required a definition of the condition of the material at the end of fatigue life and two definitions were adopted in the analysis.

Fatigue life to half initial modulus

In the construction of new pavements cemented materials where the cover the cemented material is commonly 175 mm of asphalt or more it is assumed in the Austroads Guide that when the cemented material is cracked, the cracks will readily not reflect through to the surface and the pavement has a post-cracking service life. For such pavements it is not the extent and severity of cracking in the cemented material which determines the pavement life, but commonly fatigue of the overlying asphalt. In such cases the pavement life is influenced by the magnitude of modulus of the cemented material. Jameson et al. (1992) adopted a conservative value for cemented materials fatigue life as the number of loading cycles for the modulus to decrease to half the initial value. As seen from Figure 3.2, on average this fatigue life occurred about 1/10th the number of cycles to surface cracking.


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0

2000

4000

6000

8000

10000

12000

14000

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 2.8 3.0

Ratio of cycles to cycles at 1.0 m/m2 surface cracking

Cemented materialmodulus

(MPa)

Source: Jameson et al. 1992.

Figure 3.2: Reduction in cemented materials modulus with accelerated loading

The fatigue lives so calculated and the associated predicted tensile strains data are plotted in Figure 3.3. At some chainages the cemented material had not decreased to half the initial modulus at the completion of trafficking. These chainages are marked as Censored data, whereas those that did reach half modulus are marked Observed data.


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550

1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07 Load repetitions to half the initail modulus

Tensile strain

(microstrain)

Observed data

Censored data

Fatigue relationship without constraining modulus dependency

Fatigue relationship calculated with QDOT modulusdependencyFatigue relationship calculated with Austroads modulusdependency

Source: Adapted from Jameson et al. 1992.

Figure 3.3: Fatigue relationships to half initial cemented material modulus

The Maximum Likelihood Estimation (MLE) procedure was used to derive possible fatigue relationships as this allowed both observed and censored data to be considered. Table 3.1 summarises three fatigue relationships derived assuming fatigue life was related to applied strain and cemented materials modulus as follows (Equation 5):

Ln(N) = a + bln(Strain) + cln(E) 5

where

N = number of load repetitions to half initial modulus

E = cemented materials modulus (MPa)

Strain = maximum horizontal tensile strain at the base of the cemented material (microstrain).

The coefficients derived from the MLE procedure are listed in Table 3.1. The beta values indicate how well the relationship explains the variability in results: the higher the beta value the better the fit to the data.


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Table 3.1: Models to predict fatigue life to half modulus

Constraint on modulus dependency Constant a

Strain coefficient

b

Modulus coefficient

c

Beta

Modulus unconstrained 139.7 -12.1 -7.5 9.3

Austroads modulus coefficient 130.6 -11.5 -7.0 9.2

QDoT modulus coefficient 82.5 -7.9 -3.6 7.8 Source: Jameson et al. 1992.

One key assumption of the analysis is that predicted strain and back-calculated modulus are independent variables. However, as shown in Figure 3.4, strain values are correlated with the modulus which is not unexpected as the strains values were predicted using linear elastic modelling which included cemented material modulus as an input. Hence strain and modulus data in the analysis are not independent which contravenes the assumption in the modulus unconstrained analysis and influences the reliability of the fatigue relationships derived. Hence Jameson et al. (1992) also undertook the MLE analysis with modulus dependency fixed to the Austroads and Queensland Department of Transport (QDoT) values at that time. As there was doubt about whether fatigue life varied with modulus at all, the fatigue relationship calculated with QDoT modulus dependency was favoured as it was less dependent on modulus. These three fatigue relationships are plotted in Figure 3.3 for a mean back-calculated cemented material of modulus of 10 900 MPa.

Rounding off the relationship determined using the QDoT modulus dependency, Jameson et al. (1992) derived the following relationship (Equation 6) to predict the mean fatigue life to half the initial modulus:

10

100

1000

1000 10000 100000

Cemented material modulus (MPa)

Tens

ile s

trai

n (m

icro

stra

in)


Figure 3.4: Predicted strain correlation with back-calculated modulus


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9

X = 8

0.45E35,000N

6

where

N = allowable number of repetitions of the load

= maximum horizontal tensile strain (microstrain) E = modulus of cemented material (MPa).

The use of the 8th power relationship was supported by the results of a parallel laboratory testing program (Section 3.1.1) from which a load damage exponent of 7-8 was calculated (Figure 3.1). Note that although the unconstrained modulus analysis resulted in a strain damage exponent of 12 and this provided a better fit to the data, this relationship was not favoured as the higher modulus dependency was not supported by the literature and recent laboratory testing which indicated breaking strain as the material parameter.

Based on these project findings, after allowance for the fact that the data was only obtained for one high quality material and considering the impacts of pavement thicknesses VicRoads (1993) adopted (Equation 7):

8

0.351E14,100RFN

7

where


= maximum horizontal tensile strain (microstrain) E = modulus of cemented material (MPa)

RF = reliability factor, (e.g. 1/6 for freeways).

None of the other state road agencies adopted an 8th power fatigue relationship based on the project findings. There was concern that the project had only tested one cemented material and hence the data was insufficient to change the 18th power fatigue relationship (Equation 1) then in the Austroads Guide.

Fatigue life to surface cracking

For pavements where the cover over the cemented material is less than 175 mm of asphalt it is assumed in the Austroads Guide that when the cemented material is cracked, the cracks will readily reflect through to the surface. For such pavements the fatigue life is best related to the severity and extent of surface cracking. To provide fatigue relationships for these pavement types, the fatigue life was defined as the first appearance of fine surface cracking (1.0 m/m2 cracking severity). At this amount of cracking, the modulus had reduced to about 10% of its initial value (Figure 3.2).


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The data is plotted in Figure 3.5. At some chainages the cemented material had not reached 1.0 m/m2 cracking severity at the completion of trafficking. These chainages are marked as Censored data, whereas those that did reach 1.0 m/m2 cracking severity are marked Observed data.

0

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550

1.00E+03 1.00E+04 1.00E+05 1.00E+06 1.00E+07

Load repetitions to 1.0m/m2 cracking

Tens

ile s

trai

n (m

icro

stra

in)

Observed dataCensored dataModulus unconstrainedModulus constrained to QDoT co-efficientModulus constrained to Austroads co-efficient


Figure 3.5: Fatigue relationships to 1 m/m2 surface cracking

Again, the Maximum Likelihood Estimation procedure was used to derive possible fatigue relationships as this allowed both observed and censored data to be considered. Table 3.2 summarises three fatigue relationships derived using the MLE procedure and these are plotted in Figure 3.5 for a mean modulus of 10 900 MPa.

Note again the higher strain damage exponent (14.9) of the unconstrained analysis and the associated high modulus dependency. While this relationship provided a better fit to the data, this relationship was not favoured by Jameson et al. (1992) as:

the higher modulus dependency was not supported by the literature and recent laboratory testing which indicated breaking strain as the material parameter

modulus and strain variables were not independent as discussed above.


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Table 3.2: Models to predict fatigue life to 1.0 m/m2 cracking

Constraint on modulus dependency Constant a

Strain coefficient

b

Modulus coefficient

c

Beta

Modulus unconstrained 171.3 -14.9 -9.6 11.6

Austroads modulus coefficient 138.3 -12.3 -7.0 10.9

QDoT modulus coefficient 89.4 -8.7 -3.6 8.7 Source: Jameson et al. 1992.

3.2 Further Analysis 3.2.1 Modulus Dependency The unexpected finding from the accelerated loading trial data was the very high dependency of fatigue life on modulus as seen from the modulus coefficients in Table 3.1 and Table 3.2. As illustrated in Figure 3.6, if the modulus is doubled from 5000 MPa to 10 000 MPa, the predicted fatigue life for a given strain decreases by about a factor of 100. This high dependency was not accepted by Jameson et al. 1992 and they recommended a relationship which included the modulus dependency in the QDOT relationship.

An associated unexpected finding was the high variability in surface maximum deflections and curvature (D0-D200) between experiments conducted on a 100 m long test lane with notionally identical pavement thickness and composition (Figure 3.7). The back-calculated cemented material moduli reflected this variation in measured curvature.

The reason for this curvature/modulus variation needs to be understood to reconcile the varying findings of research projects. Consequently, the results were reconsidered in this review.

0

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Tens

ile s

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n (m

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Observed dataCensored dataMaximum Likelihood Estimation E 3000 MPaMaximum Likelihood Estimation E 5000 MPaMaximum Likelihood Estimation E 10,000 MPaMaximum Likelihood Estimation E 15,000 MPa

E=5000 MPa

E=10,000 MPa

E=15,000 MPa

E=3000 MPa


Figure 3.6: Predicted dependence of fatigue on cemented material modulus


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20000

8380 8385 8390 8395 8400 8405 8410 8415 8420 8425 8430 8435 8440 8445 8450 8455 8460 8465 8470 8475 8480 8485 8490 8495

Chainage

Cem

ente

d m

ater

ial m

odul

us (M

Pa)

0.00

0.02

0.04

0.06

0.08

0.10

0.12

0.14

0.16

0.18

Initi

al 6

0 kN

Ben

kelm

an B

eam

Cur

vatu

re D

0-D

200

(mic

rom

etre

s)

Measured compressive modulus of field cores

Measured flexural modulus field beams

Back-calculated modulus

Initial 60 kN Benkelman Beam Curvatures D0-D200

Exp. 4 & 5 Exp.20Exp. 18 & 19Exp.21 Exp.6


Figure 3.7: Comparison of back-calculated moduli with measured field core and beam values

In the original analysis (Jameson et al. 1992), the Maximum Likelihood Estimation procedure was used to develop fatigue relationships as this technique enables inclusion of the censored data in the analysis. Linear regression is a simpler analysis procedure readily undertaken in spreadsheets, but is unable to consider the fact that censored data has not met the definition of fatigue failure. To assess the significance of this limitation, the half-modulus fatigue relationship obtained by linear regression was calculated for comparison with the Maximum Likelihood Estimate. The results are compared in Figure 3.8 and indicate that the linear regression line is very similar to the original Maximum Likelihood Estimation relationship. This provides confidence that analysis by linear regression (with life as the dependent variable as discussed in Section 3.1.1) is reasonable for this dataset.


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1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07 Load repetitions to half initial modulus

Tens

ile s

trai

n (m

icro

stra

in)

Observed data

Censored data

Original Maximum Likelihood Estimation- modulus as a variable

Linear regression modulus as a variable

Linear analysis without modulus and with very low life deleted

Data deleted

Figure 3.8: Reanalysis of accelerated loading data

Linear regression was used to derive a fatigue relationship without modulus as a variable and with the four data points with fatigue life less than 103 cycles deleted as follows (Equation 8):

9.18

433N

8

where


= maximum horizontal tensile strain (microstrain). Equation 8 suggests the strain dependency decreases significantly (from about 12 to about 9) when modulus is not considered in the analysis and the very low life data points are deleted.

However, due to the deletion of the modulus variable this fatigue relationship is a poor fit to the data and when the ratio of the predicted fatigue lives to the observed lives is plotted against cemented materials modulus (Figure 3.9), it is apparent the relationship over-predicts the life at high modulus and under-predicts at low modulus values.


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1000

10000

100000

0.001 0.010 0.100 1.000 10.000 100.000 1000.000

Predicted life to half modulus/observed life to half modulus

Cem

nete

d m

ater

ial m

odul

us (M

Pa)

Exp. 4, 150 mm, 40 kN Exp.6, 150 mm, 80 kNExp.8, 2x90 mm, 80 kNExp.17, 2 x 90 mm, 80 kNExp.18, 150 mm, 60 kNExp.20, 150 mm, 80 kNExp.21, 150 mm, 80 kN

Figure 3.9: Prediction error variation with cemented material modulus

It is apparent from Figure 3.10 that the twofold increase in strain due to variation in the applied load (40 kN to 80 kN) was only a minor part of the overall variation in predicted strain, with the high variation in cemented materials modulus along the test pavements the predominant source of the strain variation.

10

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1000 10000 100000

Cemented material modulus (MPa)

Tens

ile s

trai

n (m

icro

stra

in)

Exp. 4, 150 mm, 40 kNExp. 6, 150 mm, 80 kNExp.8, 2x90 mm, 80 kNExp.17, 2 x 90 mm, 80 kNExp.18, 150 mm, 60 kNExp.20, 150 mm, 80 kNExp.21, 150 mm, 80 kN

Figure 3.10: Factors influencing the variation in predicted strain


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15

Back-calculation of modulus from measured surface deflection is an inexact process and it is recognised that this is a significant source of part of the modulus variation. Nevertheless, the surface deflection did vary markedly, indicating genuine structural variation. A possible hypothesis for the large variation in modulus is that the low moduli were due to shrinkage micro-cracking during drying due to differing thermal expansion and contraction of various components (aggregates and hardened cement paste). In addition, load-induced micro-cracking prior to the commencement of accelerated loading may have occurred during the construction of the thin asphalt surfacing about a month after the 150 mm thick cemented material was placed. Given the subgrade under the cemented material was low strength (in situ CBR about seven at time of construction), delivery trucks and rollers may have induced micro-cracking in the cemented material. Consequently, it was of interest to reanalyse the data in two sets with initial moduli less and greater than 10 000 MPa, a modulus level near the mean (10 900 MPa).

The relationships are shown in Figure 3.11 and Figure 3.12 and clearly demonstrate that the higher modulus material has a higher dependence on strain than the lower modulus material. However, the question remains how should such modulus variation be addressed in routine pavement design? Does surface cracking occur at areas with initial low modulus before areas of high modulus? This issue is discussed further in Section 6.

Initial cemented material modulus less than 10,000 MPa

0

50

100

150

200

250

300

350

400

450

500

550

1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06 Load repetitions to half initial modulus

Tens

ile s

trai

n (m

icro

stra

in)

Observed data

Censored dataData deleted

Figure 3.11: Relationship for sites with initial modulus less than 10 000 MPa


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16

Initial cemented material modulus greater than 10,000 MPa

40

60

80

100

120

140

160

180

200

1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07 Load repetitions to half initial modulus

Tens

ile s

trai

n (m

icro

stra

in)

Observed data

Censored data

Figure 3.12: Relationship for sites with initial modulus greater than 10 000 MPa


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4 1997 CHANGE TO 12TH POWER RELATIONSHIP In 1994 an accelerated loading trial adjacent to the Monaro Highway, Cooma New South Wales, was carried out to assess the performance of deep-lift in situ stabilisation using cementitious binders (Jameson et al. 1995).

The gravel chosen for stabilisation was an extremely weathered granite of subbase quality similar to that on existing flexible pavements in south-eastern New South Wales. The gravel was stabilised with 5% by mass of stabilising agent consisting of 85% ground granulated slag and 15% hydrated lime.

Five experiments were carried out on stabilised pavements 250 mm, 300 mm and 360 mm thick.

The report concluded that the Austroads 18th power fatigue relationship (Equation 1) generally under-predicted the fatigue life of the trial material, whilst the Queensland Department of Main Roads (QDMR) 12th power (Equation 4) and the VicRoads 8th power (Equation 6) relationships better predicted performance.

Based on the trial findings, a literature review (Jameson 1995) and QDMR and VicRoads relationships in use in 1997, a revision to the Guide was issued by Austroads which replaced the 18th power fatigue relationships by the following QDMR 12th power relationship (Equation 9):

N = (K/)12 9and the numerator K depended on the stiffness of the material as follows:

Modulus of cemented material (MPa)

Value of K

2 000 440

3 500 350

5 000 310

10 000 260

15 000 240 In 2004 the Austroads Pavement Design Guide was published and the above K factors were used to derive the following fatigue relationship (Equation 10) suitable for any modulus in the range 2000 MPa to 10 000 MPa:

12

1910.804113,000/E

RFN

10

where


= tensile strain produced by the load (microstrain) E = cemented material modulus (MPa)

RF = reliability factor for cemented materials fatigue.


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5 RECENT RESEARCH FINDINGS

5.1 Austroads Project TT1065 Findings Following the publication of the 2004 Austroads Pavement Design Guide, further research on cemented materials fatigue has been undertaken. During 2005-2006, over 3.3 million load cycles have been applied to two cemented materials using the Accelerated Loading Facility (ALF) as part of an Austroads funded research project (TT1065 Influence of Vertical Loading on the Performance of Unbound and Cemented Materials).

The two test pavements, one comprising a crushed siltstone from Para Hills South Australia stabilised with 4% cement and the other crushed hornfels from Lysterfield Victoria stabilised with 3% cement, were constructed in late 2004 and early 2005 and then left to cure for about six months prior to accelerated loading.

5.1.1 Laboratory Testing A program of laboratory testing was conducted as part of the project (Austroads, 2008c). The testing program included both field beams cut from the test site and laboratory manufactured beams. Some of the laboratory beams were tested after 28 days moisture curing, while others were tested about 20 months after manufacture. These later beams were initially cured for three months, then removed from the fog room and stored on pallets, causing them to fully dry out. Although the samples were placed back in the fog room for a minimum 48 hours prior to testing, the wet/dry/wet cycle is thought to have influenced the material properties. For instance, the siltstone flexural modulus of the 500 days old laboratory beams was about half the value of the beams cured for 32 days.

Some of the flexural modulus results are summarised in Table 5.1.

Table 5.1: Summary for flexural moduli data

Material Field beams Laboratory manufactured beams

Age (days) Mean modulus

(MPa)

Mean relative density

(%)

Age (days) Mean modulus

(MPa)

Mean relative density

(%)

Crushed hornfels 3% cement 30 14 740 96.4 28 16 560 96.4

90+ 3 800 - 500+ 12 500 -

Crushed siltstone 4% cement 34 9 220 98.0 32 11 030 94.1

90+ 9 500 - 500+ 6 800 - Source: Austroads 2008c.

The breaking strains are summarised in Table 5.2. As seen from Table 5.2, at 28 days the hornfels field beams are more highly variable than the laboratory manufactured beams, probably due to shrinkage cracking increasing the breaking strain of some of the field beams. After further wet/dry/wet curing the breaking strain increased, but the results were more variable possible due to micro-cracking in the wet/dry/wet curing period.


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Table 5.2: Summary for breaking strain data

Material Field beams Laboratory manufactured beams

Age (days) Mean breaking

strain (microstrain)

Range breaking strains

(microstrain)

Age (days) Mean breaking

strain (microstrain)

Range breaking strains

(microstrain)

Crushed hornfels 3% cement 28 287 78-532 28 95 91-116

- - - 500+ 178 129-247

Crushed siltstone 4% cement - - - 28 339 177-500 (two beam tested)

- - - 500+ 467 295-634 Source: Austroads 2008c.

In terms of laboratory fatigue testing there were insufficient field beam results to develop a fatigue relationship. However, fatigue data on laboratory manufactured beams was available after 28 days moist curing and about 20 months wet/dry/wet curing.

The data are shown in Figure 5.1 and Figure 5.2, for the hornfels and siltstone materials respectively. As part of this review, regression analyses of the data were undertaken with the logarithm of strain as the dependent variable (refer discussion Section 3.1.1). In addition, the data was analysed in the following three ways:

using only the beams that had reduced to half modulus during the testing, without consideration of the beams that did not reduce to half initial modulus by the end of testing

using data from all beams, including those that had not reached half modulus when the loading ceased (for these beams the fatigue life was conservatively estimated as the load repetitions when the loading ceased)

using data from all beams except those with observed fatigue lives less than 1000, these were considered unreliable and below the scope of loading addressed by the Austroads design procedures.

The various fatigue relationships are plotted in Figure 5.1 and Figure 5.2. Note that the strain damage exponents were higher than those reported by Austroads (2008c) principally due to the logarithm of strain being the dependent variable in the regression analysis rather than logarithm of life (refer Section 3.1.1).

For crushed hornfels, the slope of the regression was high both for the beams tested at 28 days and 500+ days. It was noted that both the breaking strains (Table 5.2) and the tolerable strains for a given life (Figure 5.1) were significantly higher for the 500+ day old sample. This may have been due to micro-cracking of the test beam during the wet/dry/wet curing process.

The strain damage exponent for crushed siltstone was about 12, considerably lower than the crushed hornfels value (29.6). For a given strain, the crushed siltstone generally had higher fatigue life than the crushed hornfels consistent with its higher breaking strain (Table 5.2). Note however, that the siltstone fatigue results were variable and hence the 95% confidence limits on the strain damage exponent were high (95% lower limit 8, 95% higher limit 43). As seen from Table 5.2, the breaking strains of the 500+ day old siltstone were very high and suggest that the fatigue results may have been influenced by micro-cracking produced in the wet/dry/wet curing process. The presence of micro-cracking was also reinforced by the moduli (Table 5.1) at 500+ days being substantially lower than the values after 32 days. Such micro-cracking may explain


A u s t r o a d s 2 0 1 0

20

why the strain damage exponent of the siltstone was significantly lower than that calculated from the hornfels results.

Crushed hornfels

20

40

60

80

100

120

140

160

1.0E+01 1.0E+02 1.0E+03 1.0E+04 1.0E+05 1.0E+06Load repetitions to half initial modulus

Initi

al te

nsile

str

ain

(mic

rost

rain

)

Lab beams, 28 days, observed dataLab beams, 28 days censored dataLab beams, 500+ days, observed datalab beams, 500+ days, censored dataRegression 500+ days without censored dataRegression 500+ days with censored dataRegression 500+ days with censored data, lives


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5.1.2 Fatigue Under Accelerated Loading Two test pavements, one comprising a crushed siltstone from Para Hills South Australia stabilised with 4% cement and the other crushed hornfels from Lysterfield Victoria stabilised with 3% cement, were constructed in late 2004 and early 2005 and then left to cure for about six months prior to accelerated loading. The results were reported by Austroads (2008a).

As fatigue cracking was not observed on the pavement surface, fatigue relationships were derived (Austroads 2008a) based on fatigue life defined as the load repetitions to half the initial cemented material modulus. The cemented material moduli were back-calculated from the measured Falling Weight Deflectometer deflections and the measured layer thicknesses.

Two forms of data analysis were used:

The first stage was to tabulate the number of cycles to half modulus for each experiment. Knowing the applied load in each experiment, the variation in fatigue life with axle load was calculated: that is, the load damage exponents were calculated. While this analysis was instructive, the approach did not take into account variations in the pavement thicknesses, material qualities and subgrade support between experiments.

The second analysis approach addressed the variation between experiments by predicting the tensile strains applied to the cemented materials at each chainage of each experiment. Using this information, for each material, fatigue relationships were calculated relating the calculated cemented material tensile strains to the number of load repetitions to half the initial modulus.

Findings without considering variability between test sections

In terms of the first stage analysis, Table 5.3 lists the modulus reduction equations reported by Austroads (2008a), while Table 5.4 lists the calculated load damage exponents that satisfy the following equation (Equation 11):

1 Load2 Load

2 Load at Life Fatigue1 Load at Life Fatigue LDE

11

where

Load 1 = the reference load (for example the 40 kN ALF half axle load)

Load 2 = the increased load (for example the 60 kN or 80 kN ALF half axle load)

LDE = Load Damage Exponent

Fatigue Life = cycles of ALF trafficking at the given half axle load to achieve a defined percent reduction in modulus of the cemented layer. The fatigue life was determined using a logarithmic model fitted through the reduction in modulus data for each axle load.


A u s t r o a d s 2 0 1 0

22

Table 5.3: Summary of models characterising the mean reduction in modulus for each axle load

Experiment Half axle load

Model y = percent of initial modulus x = kilocycles of ALF loading

R2 n^ Mean rel. compaction

* (%)

Mean initial modulus

(MPa)

Kilocycles to 50% of initial

modulus

Hornfels material

Exp 3307 40 kN y = -0.0772Ln(x) + 0.9435 0.85 14 95.2 11 900 313

Exp 3308 and Exp 3311 # 60 kN y = -0.0931Ln(x) + 0.7463 0.95 21 95.1 6 100 14

Exp 3309 and Exp 3310 # 80 kN y = -0.0967Ln(x) + 0.6663 0.93 22 97.4 11 000 6 Siltstone material

Exp 3305 40 kN y = -0.0494Ln(x) + 0.8557 0.84 13 97.4 12 800 -

Exp 3302 50 kN y = -0.0572Ln(x) + 0.8303 0.86 7 96.3 10 500 322

Exp 3301 60 kN y = -0.0611Ln(x) + 0.7477 0.93 18 96.1 11 300 58

Exp 3304 80 kN y = -0.0701Ln(x) + 0.6498 0.94 11 96.7 11 500 8 * Mean relative compaction for the cemented base layer. # Data from experiments conducted at the same half axle load were pooled. ^ The number of cases for the regression analysis was determined by the number of FWD surveys undertaken for each experiment. Source: Austroads 2008a.

Table 5.4: Summary of Load Damage Exponents (LDEs) based on mean data at each axle load

Load ratio Extent of fatigue (percent of initial modulus)

60% 50% 40%

LDEs: hornfels material

80 kN/40 kN 5.4 5.8 6.2

60 kN/40 kN 7.1 7.6 8.2

Mean LDE 6.3 6.7 7.2

LDEs: siltstone material

80 kN/40 kN 6.4 - -

60 kN/40 kN 6.8 - -

80 kN/50 kN 7.1 7.7 -

60 kN/50 kN 8.8 9.4 -

Mean LDE 7.3 8.6 - Source: Austroads 2008a.

The critical assumption in this analysis was that each experiment had identical layer thicknesses, material qualities and subgrade support and that the only factor affecting the relative lives of the experiments was the applied load. For this reason a more rigorous analysis was undertaken by Austroads (2008a) as summarised below.

Analysis considering variability between test sections

To increase the reliability of the analysis the variability between sections was considered by calculating the strain at each chainage of each experiment from the layer moduli back-calculated from the measured Falling Weight Deflectometer surface deflections. Using this information, for each cemented material, fatigue relationships were calculated relating the calculated tensile strains to the number of load repetitions to half the initial modulus.

As seen from Figure 5.3 and Figure 5.4, when this data is plotted the results are variable and no statistically significant relationship was observed between logarithm of strain and logarithm of


A u s t r o a d s 2 0 1 0

23

fatigue life. As was the case with the Mulgrave project data (Section 3), there was a wide variation in initial modulus of each cemented material, for example:

crushed hornfels: the initial moduli varied from about 1000 MPa to 18 000 MPa, with a mean value of about 7700 MPa

crushed siltstone: the initial moduli varies from about 1000 MPa to 15 000 MPa, with a mean value of about 10 600 MPa.

As the low moduli value may have related to shrinkage micro-cracking and/or micro-cracking induced during the construction process, it was of interest to assess whether fatigue relationships could be obtained by dividing the fatigue data into two modulus groups: lower and higher than the mean modulus. For both materials, the low modulus data was too variable to deduce statistically significant fatigue relationships. However, Figure 5.5 and Figure 5.6 show the high modulus data sets and the resulting fatigue relationships. The relationships were calculated using strain as the dependent variable (refer Section 3.1.1).

Crushed hornfels: data with fatigue lives 103 to 107

20

70

120

170

220

270

320

370

420

470

1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07

Load repetitions to half modulus

Initi

al te

nsile

str

ain

(mic

rost

rain

)

3307 40 kN 3308 60kN

3309 80kN 3310 80kN

3311 60kN

Figure 5.3: Crushed hornfels fatigue results under accelerated loading: data lives 103 to 107


A u s t r o a d s 2 0 1 0

24

Crushed Siltstone: data fatigue lives 103 to 107

20

40

60

80

100

120

140

160

180

200

1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07


Initi

al te

nsile

str

ain

(mic

rost

rain

)

3301 60 kN 3302 50kN

3304 80kN 3305 40kN

Figure 5.4: Crushed siltstone fatigue results under accelerated loading: data lives 103 to 107

Crushed hornfels: data with initial cemented material modulus greater than 8,000 MPa

20

40

60

80

100

120

140

160

180

200

220

1.0E+03 1.0E+04 1.0E+05 1.0E+06


Initi

al te

nsile

str

ain

(mic

rost

rain

)

3307 40 kN

3309 80kN

3310 80kN

3311 60kN

Regression with Strain Dependent Variable

Figure 5.5: Crushed hornfels fatigue results under accelerated loading: data with initial moduli greater than 8000 MPa


A u s t r o a d s 2 0 1 0

25

Crushed siltstone : data with initail modulus greater than 10,000 MPa

40

60

80

100

120

140

1.0E+03 1.0E+04 1.0E+05 1.0E+06 1.0E+07


Initi

al te

nsile

str

ain

(mic

rost

rain

)

3301 60 kN

3302 50kN

3304 80kN

3305 40kN

Regression Strain dependent variable

Figure 5.6: Crushed siltstone fatigue results under accelerated loading: data with initial moduli greater than 10 000 MPa

5.2 Austroads TT1359 Findings In June 2007 the Austroads Pavement Structures Reference Group (PSRP) considered that the Project TT1065 findings needed to be validated by laboratory testing of a wide range of cemented materials before a change was made to the fatigue relationship in the Austroads Pavement Design Guide.

To address this need, from 2007 to 2010 fatigue testing of laboratory compacted and cured beams of a large number of materials were tested under Austroads Project TT1359 Cost-Effective Structural Treatments for Rural Highways. Findings of this research were (Gonzalez et al. 2010):

breaking strain rather than modulus is the material property influencing fatigue life the average strain damage exponent of the laboratory data was about 22, similar to the

exponent originally adopted by Austroads (1987) based on South African laboratory fatigue data (Section 2).

Table 5.5 lists the mean modulus, breaking strain and flexural fatigue relationships observed of laboratory mixed, compacted and moist cured materials.

Comparing these results to those obtained on 90+ day old field beams (Section 3.1.1 and Section 5.1.1), the laboratory mixed, compacted and moist cured materials have:

higher in modulus lower in breaking strain markedly different fatigue life dependence on strain, substantially higher strain damage

exponents.


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Table 5.5: Modulus, breaking strain and fatigue relationships for each material

Material Fatigue equation Strain damage exponent

Breaking strain (microstrain)

9 months curing

Mean modulus (MPa)

9 months curing

Weathered granite 3% log(N) = -27 x log (strain ratio) -5.45 27 185 11 500

Weathered granite 5% log(N) = -14 x log (strain ratio) -0.22 14 169 17 500

Calcrete limestone 3% log(N) = -22 x log (strain ratio) -6.2 22 286 11 500

Calcrete limestone 5% log(N) = -36 x log (strain ratio) -8.7 36 208 16 500

Basalt (Mt. Gambier) 3% log(N) = -24 x log (strain ratio) -3.89 24 216 15 000

Prior stream gravel 5% log(N) = -15 x log (strain ratio) -0.85 15 127 18 000

Modified prior stream gravel 3% log(N) = -16 x log (strain ratio) -1.77 16 126 17 500

Lean mix concrete log(N) = -21 x log (strain ratio) -2.93 21 227 (3 months curing)

36 500 (3 months curing)

Source: Gonzalez et al. 2010.


A u s t r o a d s 2 0 1 0

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6 DISCUSSION The major findings of this review of past and recent research were:

The review confirmed the current Austroads Guide procedures for predicting fatigue need major amendment as they ascribe modulus as the parameter that best differentiates the fatigue properties of materials. The recent research has validated overseas findings that breaking strain is the most appropriate material property.

Recent fatigue relationships developed from laboratory manufactured and moist cured beams indicate a high dependency on strain (Gonzalez et al. 2010). The strain damage exponents varied widely from 14 to 36, with an average value of 22. When laboratory compacted beams are dried out before testing, shrinkage of the test beams appears to have resulted in micro-cracking which reduces modulus, increases the strain at which they break and may have affected their fatigue properties.

A fatigue relationship was also derived from field beams of 3% cement treated crushed rock as part of the Mulgrave accelerated loading trial (Jameson et al. 1992). This relationship indicated a substantially lower dependence on strain than the laboratory manufactured beams: a strain damage exponent of 7-8 was calculated. The test beams were also lower in modulus than most laboratory manufactured beams and the breaking strains were very high. A possible explanation for these different properties is difference in the extent of micro-cracking. Such micro-cracking may be due to shrinkage and/or load induced during the construction of thin asphalt surfacing a month after the cemented material was placed.

Fatigue relationships have also been developed from the field performance data of accelerated loading trials (Austroads 2008a, Jameson et al. 1992). As for in-service pavements, the accelerated loading test pavements tested included areas with visible shrinkage cracking and areas of low modulus thought in part to be due to micro-cracking. As for the field beams, in general the strain damage exponents deduced from these field trials are considerably lower than those of laboratory compacted beams. The exponents calculated from the field trials were 6 to 12 for the material tested by Jameson et al. (1992) and 5 and 17 for the two materials tested more recently by Austroads (2008a).

The challenge in translating the research findings into practice is to cater for:

the significant differences in properties measured on laboratory manufactured beams and that in-service materials

the apparently high variability cemented materials properties in-service as indicated from the measured surface deflections and resulting back-calculated moduli.

One critical question to answer is whether or not low moduli areas are due to micro-cracking and whether this micro-cracking increases or decreases the load repetitions to surface cracking. This is essential in revising the Austroads structural design procedures.


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7 CONCLUSIONS The laboratory characterisation of cemented materials for road pavements has recently been completed in Austroads Project TT1359 Cost-Effective Structural Treatments for Rural Highways. This project concluded there is a need for a substantial revision to the thickness design procedures for cemented materials as provided in Austroads Guide to Pavement Technology Part 2: Pavement Structural Design (Austroads 2010). In particular there is a need to include breaking strain in the fatigue relationship to more appropriately differentiate the performance of materials. Another significant finding was that for laboratory compacted and cured beams fatigue life was more highly dependent on applied strain than the 12th power in the current Austroads fatigue relationship.

This report reviews the origin of the current fatigue relationship and previous laboratory and Australian accelerated loading trials with the objective of reconciling apparently divergent research findings and developing a research strategy for the Austroads Guide revision.

It was concluded that to revise the Austroads thickness design procedures, research is needed to:

Agree the definition of the condition of cemented materials at the end of fatigue life. Improve understanding of the reasons for the significantly different characteristics of

laboratory manufactured beams and materials in the road-bed.

Develop a method of adjusting the measured moduli of laboratory compacted and cured beams to design moduli, reflecting the initial in-service moduli of the pavement areas first to reach a terminal condition requiring maintenance.

Develop a design process to utilise fatigue data measured on laboratory compacted and cured beams. This is complicated by the fact that fatigue performance of laboratory compacted and cured cemented material beams is different from in-service pavements.

Develop a presumptive fatigue relationship applicable to the more fatigue-susceptible pavement areas.

Revise processes to design for project reliability.

In 2010/11 Austroads Project TT1664 Cemented Material Characterisation will commence with the objective of revising the structural thickness design procedures for cemented materials. It is recommended that the issues raised in this review be taken into consideration.


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REFERENCES Angell, D 1988, Technical basis for the Pavement Design Guide, report RP 1265, Pavements Branch,

Department of Main Roads, Brisbane, Qld.

Austroads 2007. Pavement design: a guide to the structural design of road pavements, errata sheet, March 2007, AP-G17/04, Sydney, NSW.

Austroads 2008a, Fatigue performance of cemented materials under accelerated loading: influence of vertical loading on the performance of unbound and cemented materials, by R Yeo, AP-T102/08, Sydney, NSW.

Austroads 2008b, Technical basis of Austroads guide to pavement technology: part 2: pavement structural design, AP-T98-08, Austroads, Sydney, NSW.

Austroads 2008c, The development and evaluation of protocols for the laboratory characterisation of cemented materials, by R Yeo, AP-T101/08, Sydney, NSW.

Austroads 2010, Guide to Pavement Technology: Part 2: pavement structural design, by G Jameson, AGPT02/10, Austroads, Sydney, NSW.

Baran, E & Aubrey, SR 1978, Preliminary report on pavement thickness design curves for Queensland conditions based on elastic layer methods, report RP531, Materials Branch, Department of Main Roads, Brisbane, Qld.

Gonzalez, A, Howard, A & de Carteret, R 2010, Cost effective structural treatments for rural highways: cemented materials, contract report, ARRB Group, Vermont South, Vic.

Jameson, GW 1995, Response of cementitious pavement materials to repeated loading, contract report CR R1949, Australian Road Research Board, Vermont South, Vic.

Jameson, GW, Dash, DM, Tharan, Y & Vertessy, NJ 1995, Performance of deep-lift in situ pavement recycling under accelerated loading: the Cooma ALF trial, 1994, APRG report no. 11, Australian Road Research Board, Vermont South, Vic.

Jameson, GW, Sharp, KG & Yeo, R 1992, Cement-treated crushed rock pavement fatigue under accelerated loading: the Mulgrave (Victoria) ALF trial, ARR 229, Australian Road Research Board, Vermont South, Vic.

Litwinowicz, A 1982, Fatigue characteristics of cement treated materials: an overview, report RP752, Materials Branch, Department of Main Roads Dept, Brisbane, Qld.

NAASRA 1987, Pavement design: a guide to the structural design of road pavements, National Association of Australian State Road Authorities, Milsons Point, NSW.

Pretorius, PC 1969, Design considerations for pavements containing soil-cement bases, PhD thesis, University of California, Berkeley.

Pretorius, PC & Monismith, CL 1972, Fatigue crack formation and propagation in pavements containing soil-cement bases, Highway Research Record, no. 407, pp. 102-15.

Symons, MG & Poli, DC 1996, Road rehabilitation by recycling: stabilisation of pavement soils from South Australia, University of South Australia, Structural Materials and Assemblies Group, The Levels, SA.


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VicRoads 1993, VicRoads guide to pavement design: supplement to Austroads Guide to the Structural Design of Road Pavements, technical bulletin no. 37, VicRoads, Kew, Vic.

Walker, RN, Paterson, WDO, Freeme, CR & Marais, CP 1977, The South African mechanistic pavement design procedure, International conference on the structural design of asphalt pavements, 4th, Ann Arbor, Michigan, University of Michigan, Ann Arbor, USA, vol. 2.

INFORMATION RETRIEVAL

Austroads, 2010, Towards the Revision of Austroads Procedures for the Design of Pavements Containing Cemented Materials, Sydney, A4, 38pp, AP-T167/10

Keywords:

fatigue, cemented materials, modulus, cracking, pavement design, breaking strain

Abstract:

The laboratory characterisation of cemented materials for road pavements has recently been completed in Austroads Project TT1359 Cost-Effective Structural Treatments for Rural Highways. It was concluded there is a need for a substantial revision to the thickness design procedures for cemented materials as provided in Austroads Guide to Pavement Technology Part 2: Pavement Structural Design. This report summarises results of past and recent laboratory and accelerated loading trials on cemented materials modulus and fatigue and proposes a research strategy that will deliver revised Austroads procedures.

a towards the revision of austroads design of pavements with cemented materials

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