Accelerated Pavement Testing

Transcription

1 Accelerated Pavement Testing 1999 International Conference October 18-20, Reno, Nevada Danish Road Institute Report Ministry of Transport - Denmark

2 Road Directorate Danish Road Institute Elisagaardsvej 5 P.O. Box 235 DK-4000 Roskilde Denmark Telephone: Telefax: vd@vd.dk web: Title: Accelerated Pavement Testing 1999 International Conference October 18-20,1999, Reno, Nevada Authors: Carsten Bredahl Nielsen, Per Ullidtz, Wei Zhang, Susanne Baltzer, Robin A. Macdonald Photo: Svend K. Larsen Dated: November 1999 Copyright: Road Directorate, All rights reserved Published by: Road Directorate, Danish Road Institute ISBN: ISSN:

3 Accelerated Pavement Testing 1999 International Conference October 18-20,1999, Reno, Nevada Carsten Bredahl Nielsen, Road Directorate Per Ullidtz, Technical University of Denmark Wei Zhang, Technical University of Denmark Susanne Baltzer, Road Directorate Robin A. Macdonald, Road Directorate Danish Road Institute Report Ministry of Transport - Denmark 3 3

4

5 Contents Preface... 6 Accelerated Testing of Permanent Deformation in Hot Rolled Asphalt Concrete Overlaid with Rut Resistant Asphalt Wearing Courses... 7 Validation of Pavement Response and Performance Models Modeling Pavement Responce and Estimating Pavement Performance

6 Preface The first paper in this report describes recent testing in the Danish Asphalt Rut Tester (DART), which has been designed and installed at the Danish Road Institute in Hot Rolled Asphalt Concrete (HRA) is used as wearing course in roughly 35% of Denmark s motorways. As there is a potential risk of permanent deformation in the HRA layer after overlaying, there is a need to adjust the wearing course to the existing pavement and to be able to test the solution. Using DART, the objective of the research is to evaluate the potential risk of permanent deformation (flow rutting) in a Danish motorway pavement by overlaying an old HRA wearing course with a new highly rut resistant wearing course. This is compared with the alternative to milling the wearing course and strengthening with a highly rut resistant binder course prior to overlaying. Structural rutting and wear rutting is not considered in this paper. The paper presents the results of accelerated testing of six slabs taken from the motorway before and after overlaying and six slabs taken from a pavement strengthened with a rut resistant binder course. The other two papers in this report describe the modeling and validation of the pavement response and performance models developed in the research work undertaken in the Danish Road Testing Machine between 1995 and As part of an international study, during , the Danish Road Institute (DRI) and the Institute for Planning of the Technical University of Denmark (DTU), constructed and instrumented a full-scale test pavement, referred to as the RTM-1 test pavement, in the Danish Road Testing Machine. During this preliminary accelerated loading test, 100,000 dual wheel load repetitions where applied, with the objective of evaluating different instruments, instrumentation and testing procedures, and to provide a preliminary pavement deterioration model for the international study. From , a second full-scale test pavement, referred to as the RTM-2 test pavement, was constructed and instrumented with similar pavement materials, structure and instrumentation, and accelerated load testing was carried out with the objective of continuing to evaluate instrumentation and testing procedures, and to complete the development of the pavement deterioration models. 150,000 load repetitions at higher load levels were applied between Based on the responses of the pavement materials and the performance of both test pavements, models for predicting subgrade and pavement deterioration and mechanistic-empirical methods for designing new and rehabilitated road pavements are proposed. 6 6

7 Accelerated Testing of Permanent Deformation in Hot Rolled Asphalt Concrete Overlaid with Rut Resistant Asphalt Wearing Courses Carsten Bredahl Nielsen, Road Directorate, Danish Road Institute Abstract Denmark has 870 km (540 mi) of motorways of which about 300 km (186 mi) have Hot Rolled Asphalt Concrete (HRA) as wearing course. After overlaying, there is a potential risk of permanent deformation in the HRA layer. Therefore, there is a need to adjust the wearing course to the existing pavement and to be able to test the solution. For this purpose the Danish Asphalt Rut Tester (DART) has been designed and installed at the Danish Road Institute in The objective of this research is to evaluate the potential risk of flow rutting in a Danish HRA motorway pavement using DART. This paper presents the results of accelerated testing of six slabs taken from the motorway before and after overlaying and six slabs taken from a pavement strengthened with a rut resistant binder course. The initial rutting increases from overlaying. This rutting is caused by a densification of the overlay and partly by a shear flow in the HRA layer. Before overlaying rutting is mainly caused by shear flow in the HRA layer. The effect of temperature on flow rutting is clearly demonstrated. The strengthening of the pavement decreases the rutting rate four times at 40 C (104 F) and five to six times at 50 C (122 F). During the summer 1998, the mean rutting level has increased 1.6 mm (0.063 in.) in-situ. This is equivalent to approximately 300 loads in DART. Based on one year of observations an estimated rutting level after 15 years of service life indicates no potential risk of rutting in the overlaid pavement 7 7

8 Introduction Background Denmark has 870 km (540 mi) of motorways of which about 300 km (186 mi) (4.2 million m 2 (45 million ft 2 )) have Hot Rolled Asphalt Concrete (HRA) as wearing course. Some HRA pavements are not very rut resistant and even after overlaying with a highly rut resistant wearing course there is a potential risk of permanent deformation in the old underlying wearing course. European Asphalt Pavement Association (EAPA) states in Asphalt - meeting the performance of Heavy Duty Pavements that if there is insufficient resistance to permanent deformation the material should be renewed at least over the thickness where the maximum shear stresses occur in the pavement. In the case of moving traffic, this is normally mm (2-3 in.) below the surface and, in the case of accelerating traffic, mm (2.5-4 in.). By applying a new 4-cm (1.5-in.) wearing course, the HRA layer would enter this zone with maximum shear stresses and there would hereby be a potential risk of flow rutting. The question is if the HRA layer should be removed or the resistance to permanent deformation is sufficient. Therefore, there is a need to adjust the new wearing course to the existing pavement and to be able to test the solution. With small rut testing devices, it seems possible to develop and test single asphalt layers, but to test overlays of motorways there is a need to test slabs with several layers taken from an existing pavement. For this purpose the Danish Asphalt Rut Tester (DART) has been designed and installed at the Danish Road Institute in Some details about DART are given below. AC Layers Tested in DART NEW Overlay: 4 cm Gap-Graded AC (1997) 4 cm HRA (1983) 5+3 cm AC (1972) 1 cm Surface Dressing (1971) 3+2 cm Soft AC ( ) 28 cm Macadam Roadbase (1956) cm Gravel Subbase (1956) Figure 1. Pavement Layers in the Danish HRA Motorway Pavement. 1 cm = 0.39 in. For this research project, a section of the motorway on the Danish Island of Funen was selected. The old HRA wearing course was going to be overlaid as part of regular maintenance. The motorway was constructed in The pavement (Figure 1, p. 8) 8 8

9 consists of cm (6-14 in.) gravel subbase, 28 cm (11 in.) macadam roadbase with a top dressing of soft asphalt concrete and 3 cm (1 in.) soft asphalt concrete wearing course. The surface was levelled out with another 2-cm (0.8 in) soft asphalt concrete after 3-5 years of traffic. The maintenance since 1961 consists of a surface dressing in 1971, a 5-cm (2-in.) asphalt concrete wearing course on top of a 3-cm (1 in.) levelling course in 1972 and a 4-cm (1.5 in.) HRA wearing course in In 1997 no serious rutting was observed in the 14 years old wearing course and it was therefore decided to overlay it with a 4-cm (1.5 in.) gap-graded asphalt concrete without milling the old wearing course. The traffic volume in one direction was in 1997 about 5,000 AADT and the traffic load in one direction about 1,000 ESALs (100 kn ( lbf)). Because of the opening of the Great Belt Bridge between the Danish Islands Zealand and Funen in the summer of 1998 both traffic volume and load has doubled. Objective The objective of this research is to evaluate the potential risk of permanent deformation (flow rutting) in a Danish motorway pavement by overlaying an old HRA wearing course with a new highly rut resistant wearing course. This is compared with the alternative to mill the wearing course and strengthen with a highly rut resistant binder course prior to overlaying. Structural rutting and wear rutting is not considered in this paper. Test Plan All samples were taken during overlay of a 6-km (4-mi) section of the motorway with a gap-graded asphalt concrete wearing course. Three of the slabs were taken before overlaying (slab 1-3) and three were taken after overlaying (slab 4-6). The alternative was carried out as a new, non-trafficked test pavement close to an asphalt plant and no milling was performed prior to paving. Six slabs were taken to study the rut resistance of the strengthened pavement (slab 7-12). All slabs were 110 cm by 140 cm (43 in. by 55 in.) in size when tested in DART. Three parameters were varied in the wheeltracking tests: A: Temperature (3 levels, see below) B: Type of loading wheel (single tyre, dual tyre) C: Type of overlay (no overlay, overlay in laboratory, overlay in-situ, strengthening). During wheel tracking tests the temperature at the surface and at the bottom of the slab were controlled and kept constant. Three sets of temperatures at the surface/ bottom were used: 40 C/20 C (104 F/68 F), 45 C/20 C (113 F/68 F) and 50 C/25 C (122 F/77 F). The loading wheel was either a 385/60 R22.5 single mounted tyre or two 315/80 R22.5 dual mounted tyres. In both cases, the loading was 45 kn (10,000 lbf) and the tyre pressure 900 kpa (130 psi). One of the slabs taken before overlaying was tested without overlaying. The two others were overlaid in the laboratory to compare with the three slabs taken after overlaying in-situ. The entire test plan is outlined in Table 1, p

10 Origin Slab No. Overlay Loading Temperature Wheel [ C] a Surface Bottom Motorway 1 In laboratory Single In laboratory Single No Single In-situ Single In-situ Single In-situ Single Test pavement 7 Strengthening Single Strengthening Single Strengthening Dual Strengthening Dual Strengthening Dual Strengthening Dual a F=1.8 C+32 Table 1. Test Plan for Wheel Tracking Slabs. Cores were drilled from the slabs after wheel tracking to determine layer thickness, air voids content and material composition. From each slab three cores were taken outside the wheel track (undisturbed area) and three cores were taken from within the wheel track to compare the situation before and after wheel tracking. Cores were drilled in-situ before overlaying to determine layer thickness, air voids content, material composition and creep properties. The cores were taken close to the position where the three slabs for wheel tracking tests were taken. Dynamic creep testing was performed using the Nottingham Asphalt Tester (NAT) on both 10-cm (4-in.) cores and 15-cm (6-in.) cores. Stiffness modulus was determined on 10-cm (4-in.) cores using NAT. Drilled cores from the motorway after overlaying were used to determine creep and stiffness properties of the new overlay. Experimental Procedures Sampling of Slabs Slabs were cut from the motorway and from the non-trafficked test pavement close to an asphalt plant. Slabs were cut to form rectangular specimens of 110 ± 5 cm by 140 ± 5 cm (43 ± 2 in. by 55 ± 2 in.). The thickness of the slabs was equal to the thickness of all the bound layers in the pavement. The slabs on the motorway were sampled in a straight line in the heavy lane of the two-lane motorway. They were sampled so that loading in the wheel-tracking test was applied in the same direction as traffic on the motorway. The distance from the centre of the slab to the edge of the pavement was approximately 100 cm (40 in.) to avoid cracked areas close to the edge of the pavement. The distance between slabs in traffic direction was 200 cm (80 in.) to allow cores to be drilled in between slabs. After cutting, slabs were left in place and a 20-cm (8-in.) groove was thoroughly cleaned up along all four sides in full depth of the slab. The surface of the slab was cleaned thoroughly and a special designed steel box with 10 10

11 internal adjustable plane steel plates along the sides were placed over the specimen. The surface of the slab was in full contact with the supporting steel plate (the bottom of the steel box) inside the steel box before spanning the adjustable steel plates along the sides of the slab. When the slab was sufficiently spanned, it was turned upside down. In this way, three slabs were cut before overlaying. Another three slabs were cut during overlaying at the end of the first day stop. The slabs were cut after cooling of the wearing course approximately 25 m (80 ft) before the day stop within a normal roller pattern. The next day repairs were made before continuing overlaying. In this way, repairs were avoided in the new wearing course. Therefore, sampling of slabs during overlaying was controlled of the position of the day stop and it was not possible to take samples before and after overlaying close to each other. The distance between the two positions was approximately 1.3 km (0.8 mi). Preparation of the Slabs at the Danish Road Institute The slabs were transported upside down in a horizontal position on a truck. Upon arrival at the Danish Road Institute, the slabs were inspected for cracks and deformations occurred during handling and transport. A weak Portland cement mortar was used to level the underneath of the slabs out to ensure a parallel, plane support during testing. The slab was then turned upside down by placing a 30-mm (1-in.) supporting steel plate on top of the levelling layer, spanned to the steel box. On two slabs, an overlay was compacted in laboratory. The overlay was mixed at an asphalt plant as a special production. The asphalt was transported to the Danish Road Institute on a small truck with an insulated cover during transport. The overlay was rollercompacted in a concrete mould with steel sides on a 30-mm (1-in.) steel-plate. The bottom of the mould was vertically moved to obtain the prescribed slab thickness. A Bomag BW90 S double drum vibratory compactor with a total weight of 1,300 kg (1.4 T) was used for compaction. The slabs were placed into the test mould with overall dimensions of 120 cm by 150 cm (47 in. by 59 in.), leaving a 5-cm (2-in.) gap to eliminate minor inaccuracies from cutting the slab. The gap was filled with carefully compacted 0/8 mm (0/0.31 in.) gravel and at the upper 5 cm a weak 0/8 mm (0/0.31 in.) Portland cement mortar

12 The Danish Asphalt Rut Tester (DART) Wheel tracking tests were performed in DART at the Danish Road Institute. The draft European standard for Wheel Tracking pren is adapted to DART. A detailed description of DART is given in (1). The device simulates a rolling load with side wander. The estimated number of repeated loads is 24,000 a day (bi-directional) and 12,000 a day (unidirectional). Data for the device are in brief: Stationary heavy vehicle simulator with linear travel 0-5 km/h (0-3 mi/h) Wheel load up to 65 kn (14,500 lbf.) (corresponding to 130 kn (29,000 lbf.) axle load) Random normally distributed wander (± 20 cm (± 8 in.) from centre travel) Full-size standard lorry tyre, single or dual wheel configurations Automatic tyre pressure control Test sample 120 cm by 150 cm (47 in. by 59 in.), thickness 5-25 cm (1-10 in.) Air temperature control cabinet, C ( F) Temperature control radiator underneath the slab, C ( F) Separate pre-conditioning cabinet for slab Automatic control, allowing temperature conditioning and load applications overnight Automatic rutting and macro texture measurements with precision laser profilometer Manual control and tyre contact stress calibration Figure 2. Outline of the Danish Asphalt Rut Tester (DART) 1 Loading Wheel, 2 Loading Cart, 3 Loading Arm, 4 Hydraulic Cylinder, 5 Pneumatic Cylinder, 6 Test Mould, 7 End platform with Roller, 8 End Platform with Tyre Contact Stress Calibration (Home Position), 9 Profilometer Cart, 10 Laser Profilometer

13 DART is built into a conditioning cabinet of which air temperature can be controlled from 25 C to 60 C (77 F to 140 F). The slab is mounted in a test mould equipped with an adjustable, supporting plate. The temperature of the underneath of the slab can separately be controlled from 25 C to 60 C (77 F to 140 F). During testing temperature is kept constant at the surface and underneath the slab. Therefore, it is possible to obtain a constant temperature gradient in the slab. DART consists of a large steel frame 6.6 m (22 ft) long, 3.0 m (10 ft) wide and 3.9 m (13 ft) high (Figure 2, p. 12). The loading wheel is mounted in a cart travelling on a top rail driven by belt and electric motor. On the cart, the loading wheel has a possible wander of ± 20 cm (± 8 in.). The load is applied by means of a hydraulic cylinder at one end of the cart. Below the loading wheel are three platforms that can be moved vertically to adjust to the wheel size. The test mould is mounted on the middle platform and the two end platforms are adjusted to the same level. The wheel accelerates at the end platforms to a maximum speed of 5 km/h (3 mi/h). The loading wheel is equipped with a full-size lorry tyre either in single or dual configurations. The user specifies tyre pressures up to 1.0 MPa (145 psi), and the specified pressure is automatically kept constant by compressed air during testing. Wheel load is only applied on the slab to avoid oscillations from the loading wheel as it passes the edge of the test mould. The load is applied in 0.2 seconds to a maximum of 65 kn (14,500 lbf) and is relieved before the wheel leaves the slab. This gives a travel of 64 cm (25 in.) with constant load and speed in the middle of the slab. The wander of the loading wheel is randomly selected and normally distributed The user specifies mean position and standard deviation. In one position, ten loads are applied, and then, before the wheel wanders to the next position, the wheel is randomly rotated at one end platform. A laser mounted on a small cart travelling on top rail measures rutting. The laser is automatically positioned in all directions. The laser measures five cross and three longitudinal profiles. The user specifies the position of these profiles. The laser measures one point every mm with a vertical accuracy of 0.1 mm (0.004 in.). Measuring is performed while the loading wheel is resting at the opposite end platform. Experimental Results Material Composition After completion of wheel tracking tests, six cores were drilled in each slab. Three cores were taken outside the wheel track (undisturbed area) and three cores were taken from within the wheel track to compare the situation before and after wheel tracking. It is assumed that cores from outside the wheel track represents the situation in-situ when the slabs were sampled. Thickness and density of each layer was determined for all cores from each slab. Material composition in each layer was only determined for cores from one slab for each of the following groups: Slabs sampled prior to overlaying (slab 1), slabs sampled after overlaying (slab 5) and slabs from test pavement (slab 7). For each group, grading curves are given for all layers in Figure 3, p 14. Bitumen grades were characterised by penetration at 25 C (77 F) (Pen.) and softening point (R&B) of recovered binder from material samples. The 13 13

14 average layer thickness and mix composition is given for each group in Table 2, p. 15. Voids in mineral aggregate (VMA), voids content and penetration index (PI) are calculated and also given in Table 2, p. 15. Wheel Tracking Tests During wheel tracking tests, the slabs were conditioned at a constant temperature at the surface and at the bottom of the slabs. These temperatures were controlled within ±1 C (±1.8 F) and sampled 4 times an hour. Temperatures within the slabs were not measured during testing. Measurement of temperature profiles in two slabs during conditioning prior to testing (slab 7 and 8) indicates that temperature remains constant with a linear gradient after hours of conditioning. Therefore, all slabs were conditioned at least 24 hours prior to testing. From the distance between the surface of the slab and the steel supporting plate (with known temperatures) temperatures in the different layers in the slab are calculated assuming a linear temperature gradient. For slab 4-6, this distance was 240 mm (9.4 in.) and for all other slabs it was 220 mm (8.7 in.). The average layer thicknesses in Table 2, p. 15 are used to calculate the temperature in the middle of each layer. The results are given in Table 4, p. 18. The total number of loads was not fixed but was adapted to the rut resistance of the slab to obtain as many loads as possible within the measuring range of the laser profilometer. The total number of loads is given in Table 3, p. 17 for each slab. 100 AC gap (slab 1-2) HRA (slab 1-3) Percent passing ,01 0, Sieve size [mm] AC con (slab 1-3) AC lev (slab 1-3) AC gap (slab 4-6) HRA (slab 4-6) AC con (slab 4-6) AC lev (slab 4-6) AC gap (slab 7-12) AC bind (slab 7-12) AC lev (slab 7-12) Figure 3. Grading Curves for All Layers in the Slabs Used for Wheel-Tracking Tests 1 mm=0.039 in

15 Layer Year Thickness [mm] a Bit. Cont. [%m] VMA Voids Cont. R&B [ C] b Pen. [dmm] a PI Slab 1-3 AC gap-graded % 9.6% HRA % 2.4% AC con.-graded % 4.5% AC levelling course % 5.9% Slab 4-6 AC gap-graded % 11.2% HRA % 1.6% AC con.-graded % 6.6% AC levelling course % 4.1% Slab 7-12 AC gap-graded % 9.4% Binder course % 5.6% AC base course % 6.5% a 1 mm = in. 1 dmm = in. b F = 1.8 C+32 Table 2. Composition of Slabs Tested in DART. Calculations of Profiles and Rutting Rates For any specified number of loads, five cross profiles and one longitudinal profile were measured with the laser profilometer. Cross profiles were measured at the cross centre line of the slab and ± 100 mm (4 in.) and ± 200 mm (8 in.) from this. The longitudinal profile was measured at the longitudinal centre line of the slab. Values of the vertical distance from a horizontal reference plane to the surface of the slab were measured. Each cross profile consists of 1,500 values and each longitudinal profile of 1,000 values with a onemm horizontal distance. First all profile measurements were filtered to exclude out-of-range data caused by missing reflections from steep changes in surface texture. It is assumed that all five cross profiles are identical. Each value in a profile was compared with the five values before and after in the same profile and the five values before and after in the other four profiles. In this way, each value was compared with the 54 closest values. Values differing more than 20 mm (0.8 in.) from the minimum of the other 54 values were rejected. These values are likely to be out of range. The same comparisons were made for the longitudinal profile but each value was only compared to the closest 10 values in the same profile. The mean cross profile and the mean longitudinal profile was calculated for all specified number of loads from the filtered profiles. Each mean value is an average over the same values as used for filtering. The initial mean cross profile and the initial mean longitudinal profile was calculated from the first measurement sequence performed prior to loading. The rut depth cross profile and the rut depth longitudinal profile was calculated for all specified number of loads by subtracting the initial mean profile from all other mean profiles. The rut depth profiles after 20,000 loads are given for slab 3, 5, 7 and 9 in Figure 4, p. 16. The rut depth, P N, for each specified number of 15 15

16 loads was calculated from the rut depth profiles as an average over the central 200 mm (8 in.) of the rut depth cross profile and the central 400 mm (16 in.) of the rut depth longitudinal profile Rut depth [mm] Slab 9 Slab 7 Slab 5 Slab Cross position [mm] Figure 4. Rut depth profiles after 20,000 Loads for slab 3 (Reference, No Overlay), Slac 5 (In Situ Overlay), Slab 7 (Strengthened Pavement, Single Wheel) and Slab 9 (Strengthened Pavement, Dual Wheel). 1 mm= in. The rut depth, P N, at N loads is estimated by the power function: P N = A N b (1) By linear regression, the constants A and b in the power model are derived from experimental data. In Figure 5, p. 17 the measured rut depth is plotted against the number of loads for all slabs and the power models are given. The slope or rutting rate, P, is calculated at 20,000 loads from differentiation of the rut function. P 20,000 = A b 20,000 (b-1) (2) 16 16

17 Rut depth [mm] Number of loads Slab 1 Slab 2 Slab 3 Slab 4 Slab 5 Slab 6 Slab 7 Slab 8 Slab 9 Slab 10 Slab 11 Slab 12 Figure 5. Measured Rut Depth against the Number of Loads and Regression Models for All Slabs. 1 mm=0.039 in. Slab No Loads Constants in power model Rutting rate A b P 20,000 [µm/load] a 1 27, , , , , , , , , , , , a 1µm = in. Table 3. Results of the Wheel Tracking Tests for Each Slab

18 Layer Testing Temperature [ C] a Slab 1 AC gap-graded 38 8% -1% HRA 34 9% -1% AC con.-graded 30 4% 0% AC levelling course 26 14% -1% Slab 2 AC gap-graded 38 24% 2% HRA 34 12% -1% AC con.-graded 30 6% 0% AC levelling course 26-27% -2% Slab 3 HRA 38 46% 1% AC con.-graded 34-3% 2% AC levelling course 30 29% 0% Slab 4 AC gap-graded 48 39% 9% HRA 44 57% 2% AC con.-graded 39 10% 1% AC levelling course 36 26% 2% Slab 5 AC gap-graded 38 29% 7% HRA 35 56% 1% AC con.-graded 32 10% 1% AC levelling course 29-13% 2% Slab 6 AC gap-graded 43 32% 5% HRA 39 19% 1% AC con.-graded 34 10% 0% AC levelling course 31 0% 0% Slab 7 AC gap-graded 39 11% 5% Binder course 34 0% 2% AC base course 28 0% 1% Slab 8 AC gap-graded 48 20% 8% Binder course 42 12% 5% AC base course 34 5% 1% Slab 9 AC gap-graded 39 10% 5% Binder course 34 4% 2% AC base course 28 0% 1% Slab 10 AC gap-graded 39 14% 3% Binder course 34 0% 1% AC base course 28-1% 1% Slab 11 AC gap-graded 48 14% 7% Binder course 42 11% 3% AC base course 34 3% 2% Slab 12 AC gap-graded 48 9% 5% a F = 1.8 C+32 Proportional Binder course 42 4% 2% AC base course 34 13% 2% Rut Proportional Densification Table 4. Results of the Wheel Tracking Tests for each Layer

19 Results are given in Table 3, p. 17. The proportional rut in each layer are calculated as the difference in layer thickness between layers outside and within the wheel track in relation to the thickness outside the wheel track. The proportional densification are calculated as the difference in layer density between layers outside and within the wheel track in relation to the density outside the wheel track. Results are given in Table 4, p. 18. Dynamic Creep Tests Dynamic creep tests were performed using the Nottingham Asphalt Tester (NAT) on cores drilled from the pavement prior to overlaying. Only the HRA layer and the underlying AC layer were tested. Two groups of cores were tested: The first group having a diameter of 10 cm (4 in.) and the second group having a diameter of 15 cm (6 in.). In each group eight cores were tested. Four cores were cut to have a 30-mm (1.2-in.) HRA and a 40-mm (1.6-in.) AC layer each and were tested as double layers. The other four cores were separated into a 30-mm (1.2-in.) HRA layer and a 40-mm (1.6 in.) AC layer and were tested as single layers. Testing was performed according to FAS Method (1997) at 40 C (104 F) for both 10-cm (4-in.) cores (unconfined) and 15-cm (6-in.) cores (confined). In addition, dynamic creep tests were performed on four 10-cm (4-in.) cores drilled from the overlay on the motorway. Thickness and material composition of each layer was determined prior to testing. The constants A and b in the power model are derived from experimental data in the same way as for the slabs, and the slope or rutting rate is calculated at 3,000 loads from differentiation of the power function. The results are given in Tables 5 and 6, p. 19. In Table 6, p. 20 an estimate of the rutting rate for double layers is calculated by adding creep measured during testing of single layers. 10-cm cores AC gap-graded % 12.1% HRA % 2.4% AC con.-graded % 4.8% cm cores HRA % 3.1% AC con.-graded % 3.7% a 1 mm = in. 1 dmm = in. b F = 1.8 C+32 Layer Year Thickness Bit. VMA Void R&B Pen. PI Cont. Cont. [ C] b [dmm] a [mm] a [%] Table 5. Composition of Cores for Dynamic Creep Tests (NAT)

20 Layer Constants in power model Rutting rate Material 10-cm cores AC gap-graded HRA AC HRA + AC a HRA and AC b cm cores HRA AC HRA + AC a HRA and AC b a Estimated by adding creep measured during testing of single layers b Tested as double layers c 1µm = in. Thickness [mm] A b P 3,000 [µm/load] c Table 6. Results of Dynamic Creep Tests at 40 C (104 F). Evaluation of Experimental Data The original pavement prior to overlaying serves as a reference for the testing of permanent deformation. Compared to other asphalt pavements with HRA the total thickness of the asphalt layers is rather low. The original pavement is one of the oldest motorway pavements in Denmark and it is designed with an unbound granular base and repaired with several overlays. The HRA overlay has a low void content with a reasonable stiffness of binder and mix. The underlying AC layers have a normal void content and a reasonable stiffness. The lower AC layer, which was originally soft, has become very hard, and all slabs disintegrated in this layer when they were lifted out from the pavement. An overall evaluation of material data leads to the conclusion that the slabs taken before and after overlaying is expected to have approximately the same structural properties including all layers from the HRA layer and below. The following discussion is based on this assumption. Using slab 3 as a reference, the following points will be discussed: How does the overlay affect flow rutting? How does temperature affect flow rutting after overlaying? How much is flow rutting changed by strengthening? How does the tyre type affect flow rutting in the strengthened pavement? Does densification or shear flow cause flow rutting? Is it possible to estimate wheel-tracking results from dynamic creep tests? Various authors have suggested models to predict rut depth in bituminous pavements. Verstraeten (3) gives an overview of these models in a comprehensive presentation of all aspects of highly rut-resistant bituminous mixes. In the present paper, though, only empirical models will be used. The discussion is concentrated on the measured rutting rate at 20,000 loads and the calculated rutting and densification in each layer. It should 20 20

21 be noticed that the estimated rutting in each layer is not very accurate as it is based on measurement of layer thicknesses on drilled cores with inclined boundaries between layers. The Effect of Overlay on Flow Rutting Slab 1, 2 and 5 with overlays are all tested at the same surface temperature as the reference slab without overlay. The average rutting rate is reduced by 15% compared to the reference. The effect of overlaying is partly explained by the reduction in temperature of the HRA layer from 38 C (100 F) in the reference to 34 C (93 F) in the overlaid slabs. There will also be a similar reduction in temperature after overlaying the HRA layer in-situ. The initial rutting increases from overlaying in-situ. This rutting is caused by a densification of the overlay and partly by a shear flow (rutting without densification) in the HRA layer. Before overlaying rutting is mainly caused by shear flow in the HRA layer. When comparing rutting levels of slab 5 and the reference it should be noticed that the number of loads is more than doubled in slab 5. The reduction of the rutting rate is 25% compared to slab 5 with in-situ overlay and only 9% in average compared to slab 1 and 2 with overlay compacted in laboratory. The difference between overlaying in-situ and laboratory is explained by a difference in grading curve of the overlay (Figure 3, p. 14). Although the overlay used in laboratory was intended to have the same composition as that in-situ, this is clearly not the case. The mix used in laboratory is extremely gap-graded and was observed to have a tendency towards bleeding. Compaction of the overlay was poorer in slab 2 than in slab 1 and the rut in the overlay in slab 2 is also observed to be larger. Humps in the cross section profile close to the boundary of the test mould indicate that shear flow is the main cause for rutting in the slabs overlaid in laboratory. The shear flow in the laboratory compacted overlay is much larger than in the in-situ compacted overlay (Figure 6, p. 23). Overlaying in laboratory and overlaying in-situ is therefore clearly different in this research. The results indicate that it is important to use exactly the same mix composition and the same level of compaction. The Effect of Temperature on Flow Rutting after Overlaying The effect of temperature on flow rutting is clearly demonstrated by comparing the rutting rate of slabs 4-6. When the temperature is increased by 10 C (18 F) from 38 C (100 F) to 48 C (118 F) in the overlay, the rutting rate is tripled. This effect is much larger than the effect of overlaying. Initial rutting in the overlay is of the same order of magnitude at all three temperature levels. When comparing the densification and rut level in the different layers it should be noticed that the number of loads is different. If the parameters are normalised to the same number of loads (assuming that the densification and rut are linear with the number of loads) the same effect of increasing temperature is observed in all layers. Rut and densification increase dramatically with temperature. In average, it seems that about one half of the rutting occurs in the HRA layer, one third in the overlay and the rest in the underlying AC layer

22 The Effect of Strengthening on Flow Rutting The alternative to overlay is to mill the HRA-layer and strengthen with a rut resistant binder course prior to overlaying. It would have been interesting to see the effect of this repair in-situ but for practical reasons, it was not possible. Instead, a nontrafficked test pavement was constructed. From Figure 3, p. 14 and Table 2, p. 15 it is seen that mix composition and compaction level of the overlay is similar to the in-situ overlay. It is therefore assumed that the strengthened pavement is an acceptable substitute for the in-situ repair. The rutting rate is decreased by a factor of four compared to the overlay without strengthening when the surface temperature is 40 C (104 F) (comparing slab 5 and 7). Almost all rutting occurs in the overlay. At a surface temperature of 50 C (122 F), the rutting rate is decreased 5-6 times (comparing slab 4 and 8). At this temperature rutting and densification is also observed in the bindercourse. It is obvious that the strengthening has a marked effect on the rutting resistance of the pavement. From the grading curve (Figure 3, p. 14) it can be seen that the binder course is gap-graded and therefore, the rut resistance of the binder course is achieved by the stone to stone contact. The Effect of Tyre Type on Flow Rutting In the wheel-tracking tests of slab 8-12, the loading wheel was mounted with dual tyres (315/80 R 22,5) instead of a single tyre (385/65 R 22,5) used for testing all the other slabs. The load and tyre pressure was the same for the two types of tyres. Roughly, the effect therefore is a reduced contact area and loading time. The contact area is 290 mm (11 in.) wide for the single tyre and 235 mm (9 in.) for each of the dual tyres. The distance between dual tyres contact areas is 120 mm (4.7 in.). Assuming a rectangular contact area this gives a reduction in loading time from 120 ms to 75 ms at the actual wheel-tracking speed. This increases pavement stiffness and hereby decreases rutting. In practice, contact stresses are expected to vary within the contact area and be different for the two types of tyres as observed by Krarup (3). In fact, shifting from a single tyre to dual tyres decreases rutting rates. The average rutting rate for slab 9 and 10 tested with dual tyres is less than half the rutting rate of slab 7 tested with a single tyre. When the temperature is increased (comparing slab 11 and 12 with slab 8) the rutting rate using dual tyres is less than one fourth of the rutting rate using a single tyre. This indicates that the significance of the contact stress distribution increases with temperature. Densification and Shear Flow All rutting rates are calculated from the measured profiles. In these profiles rutting is defined as the reduction in the thickness of the slab caused by repeated passes of the loaded wheel. The profiles do not include the humps to the sides of the wheel track in the slab. The humps can be included if the wheel-track depth is measured as the distance from a straightedge connecting the tops of the humps to the surface of slab. The wheel-track depth is in most cases roughly the double of the rut depth, but this varies with the properties of the mix. In some cases, the wheel-track depth is meaningless if the humps are not regular but have a local maximum. This causes an overestimated wheel-track depth as the straightedge is placed at this maximum. Therefore, it makes more sense to calculate the volume of the track and the humps from the measured cross profiles. A local maximum will only have a minor influence on the volume of 22 22

23 the hump. The volume is estimated from the area below and above the initial profile. The ratio between the area above and below the initial profile is a useful measure of the ratio between densification and shear flow. Pure densification would cause a ratio close to zero, as no humps would be formed. Pure shear flow on the other hand would cause a ratio close to one as no densification would take place, and all material pushed down in the wheel track would be pushed up as humps at the sides. Therefore, it makes sense to define this ratio as the shear flow percentage. The cross profiles are measured as a function of the number of loads and consequently it is possible to analyse the flow rutting mechanisms as function of the number of loads. In Figure 4, the shear flow percentage is given as function of the number of loads for all slabs tested. It should be noticed that the shear flow percentage only gives the ratio between shear flow and densification. It does not say anything about the magnitude of rutting or rutting rate. 100% 80% Slab 1 Slab 2 Slab 3 Shear flow 60% 40% 20% Slab 4 Slab 5 Slab 6 Slab 7 Slab 8 Slab 9 0% Number of loads Slab 10 Slab 11 Slab 12 Figure 6. Shear Flow Percentage against the Number of Loads for All Slabs From Figure 6, p. 23 it is seen that all slabs from the motorway pavement have a shear flow between 80% and 90% at the end of testing. The shear flow percentage in the strengthened pavement is considerably less. In general, pavements with a low rutting rate have a low shear flow percentage. It is clear that temperature affects the shear flow percentage but the shear flow percentage also depends on the mix-type. Slab 2 was found to have a relative high rutting rate compared to slab 1 and 5 and this is apparently caused by an unstable overlay with pure shear flow. It is also noticed that the pavement overlaid in laboratory develops shear flow quicker than the reference pavement without overlay. The pavement overlaid in-situ develops shear flow slower than the reference pavement. The shear flow percentage in the strengthened pavement is larger at a surface temperature of 50 C (122 F) than at 40 C (104 F). This explains the higher rutting rate at higher temperatures. The dual wheel configuration leads to a 23 23

24 lower shear flow percentage than the single wheel configuration and the effect of temperature is seen here. The only exception is slab 9 tested with a dual wheel configuration at 40 C (104 F) which was expected to be identical with slab 10. Although the two slabs have the same rutting rate, the initial shear flow percentage in slab 9 is much larger and is decreasing with the number of loads. This effect can not be explained in a simple way. Estimating Wheel Tracking Results from Dynamic Creep Tests The dynamic creep test shows a very large difference between the estimated rutting rates for 10-cm (4-in.) cores (unconfined) and 15-cm (6-in.) cores (confined). The unconfined rutting rate is more than 20 times that of the confined test. The tests also show that the rutting rate for two layers can not be estimated by simply adding the creep for each of the layers. For the unconfined specimens, the rutting rate estimated from adding the creep of each layer is only half of the rutting rate of the double layer specimens. For the confined cores it is only one fourth. Therefore, it seems not possible to estimate rutting rates for multi-layer specimens in a simple way. Wheel tracking tests and dynamic creep tests are not directly comparable, as loading conditions and temperature are different. The load in dynamic creep tests are 100 kpa (14.5 psi) compared to a 900 kpa (130.6 psi) load in the wheel tracking tests. The dynamic creep tests indicates that it is only possible to compare rutting in wheel tracking tests with dynamic creep measured on double layer specimens because no simple way to add rutting from single layers was found. In the wheel tracking tests, it was found that most rutting occurred in the HRA and the underlying AC layer. Therefore, it might be possible to estimate wheel tracking results for the reference pavement (slab 3) from dynamic creep results. The result is given in Figure 7, p. 24. The correlation coefficient between creep tests and wheel tracking tests is quite good. The usefulness of this relation, though, is doubtful as it is only valid for this combination of layers and the estimated constants can not be used for other pavements. Dynamic creep rate [um/load] y = 0,4424x + 0,6826 R 2 = 0,944 y = 0,0221x + 0,0369 R 2 = 0, Wheel tracking rate [um/load] 10-cm (4-in.) cores 15-cm (6-in.) cores Figure 7. Correlation between Wheel-Tracking Rate and Dynamic Creep Rate (NAT). 1 mm= in

25 Implementation Strategy The objective of the research project is to evaluate the potential risk of permanent deformation (flow rutting) in a motorway pavement by overlaying an old wearing course with a new highly rut resistant wearing course. The results from this research project will be implemented in future overlaying strategies in larger projects. Initial testing of the in-situ pavement combined with overlaying in laboratory and testing in DART is meant to lead to the most suitable and economic overlaying strategy for the actual pavement. To be able to do this the following tasks has to be solved: Better control with overlaying in laboratory Mathematical models to estimate rutting in-situ from wheel tracking results, actual traffic and temperatures in-situ. Correlation of wheel tracking results with in-situ rutting. Rutting in-situ has been measured with the Danish Profilograph before and after overlaying. The motorway was overlaid in June The mean rutting level for both wheel tracks in the heavy lane was 11 mm (0.43 in.) in July 1996, one year before overlaying. This was reduced by overlaying to 1.9 mm (0.074 in.) in July 1997 and two month later in September 1997 a small increase to 2.5 mm (0.098 in.) was observed. The latest measurement was performed in October after the summer of The mean rutting level has now one year after overlaying increased to 4.1 mm (0.161 in.). The mean rutting level on the motorway is given in Figure 8, p. 25 between kilometres 133 and 135. Slabs 1-3 were taken in kilometre and slab 4-6 in kilometre as indicated in Figure 8, p Slabs 4-6 Slabs 1-3 Rut depth [mm] Kilometre Figure 8. Rutting Measured with the Danish Profilograph on the Motorway where Slabs 1-6 Were Taken 1 mm=0.039 in

26 During the summer 1998 the mean rutting level has increased 1.6 mm (0.063 in.). Comparing with the wheel tracking test of the in-situ overlaid slab 5 tested at a surface temperature of 40 C (104 F) this increase is equivalent to approximately 300 loads. This means that in-situ there has only been 300 loads equivalent to the loading conditions and temperature in the wheel-tracking test. Assuming this equivalent number of loads, the other wheel tracking results at 40 C (104 F) (slab 3 and slab 7) can be used to estimate the rutting level after the first summer. The rutting level is estimated to 1.0 mm (0.039 in.) in the pavement without overlay and 1.6 mm (0.063 in.) in the strengthened pavement. This is not surprising, as the initial rut will take place in the overlay. Assuming the same yearly equivalent number of loads, the estimated rutting levels after 15 years of service life would be 7 mm (0.28 in.) with overlay, 6 mm (0.24 in.) without overlay and 4 mm (0.16 in.) with strengthening. As a rutting level up to 20 mm (0.79 in.) would be acceptable, there is no need to overlay or strengthen the pavement if only rutting is considered. However, for other reasons an overlay was advantageous. Based on one year of observation in-situ there is no potential risk of rutting by overlaying of this motorway pavement. The estimate is of course very uncertain as the equivalent number of loads might increase due to an increase in traffic or the number of days with hot weather. In 15 years, a yearly equivalent number of loads of 2,000 would cause an estimated rutting level of 20 mm (0.79 in.) in the overlaid pavement, 23 mm (0.91 in.) in the original pavement and 7 mm (0.28 in.) in the strengthened pavement. This indicates that the strengthening only is advantageous if the traffic or the number of hot days is expected to increase drastically. On-going research projects are trying to improve the estimation of equivalent loads from correlation of wheel tracking tests in DART with observations of rutting in-situ. In this research, other factors like the durability of the pavement are taken into consideration to optimise the service life of the pavement. With time, this will lead to a more safe estimation of rutting resistance from wheel-tracking tests. Summary and Conclusions This paper presents the results of accelerated testing of permanent deformation (flow rutting) in a Danish HRA motorway pavement by overlaying an old wearing course with a new highly rut resistant wearing course. This is compared with the alternative to mill the wearing course and strengthen with a highly rut resistant binder course prior to overlaying. Structural rutting and wear rutting is not considered in this paper. The objective is to evaluate the potential risk of flow rutting in a motorway pavement using the Danish Asphalt Rut Tester (DART) the new full-scale wheel tracking machine at the Danish Road Institute. From the results of this investigation the following conclusions can be drawn: The initial rutting increases from overlaying. This rutting is caused by a densification of the overlay and partly by a shear flow in the HRA layer. Before overlaying rutting is mainly caused by shear flow in the HRA layer. Overlaying in laboratory and overlaying in-situ are clearly different in this research. The effect of temperature on flow rutting is clearly demonstrated. When the surface temperature is increased by 10 C (18 F) from 40 C (104 F) to 50 C (122 F) in the overlay, the rutting rate is tripled. In average, about one half of the rutting occurs in 26 26

27 the HRA layer, one third in the overlay and the rest in the underlying AC layer. The strengthening of the pavement decreases the rutting rate four times at a surface temperature of 40 C (104 F) and five to six times at 50 C (122 F). A further decrease is observed by shifting from a single tyre to dual tyres. From wheel tracking tests, a measure of shear flow is defined. The shear flow in the strengthened pavement is considerably less than in the motorway pavement. In general, pavements with a low rutting rate have a low shear flow. The dual wheel configuration causes less shear flow than the single wheel configuration. A good correlation between wheel-tracking tests and dynamic creep tests is observed for the pavement without overlay. The usefulness of this relation is doubtful, as it is only valid for a specific combination of layers. It has not been possible to estimate rutting rates for multi-layer specimens from dynamic creep testing of single layers in a simple way. During the summer 1998, the mean rutting level has increased 1.6 mm (0.063 in.) in-situ. This is equivalent to approximately 300 loads at the loading conditions and a surface temperature of 40 C (104 F) in the wheel tracking test. Based on one year of observations the estimated rutting level after 15 years of service life would be 7 mm (0.28 in.) which indicates no potential risk of rutting by overlaying of this motorway pavement. References (1) Nielsen, C.B. New full-scale Danish Rut Tester. Nordic Road & Transport Research Vol. 8, No. 3, Dec. 1996, pp (2) Krarup, J. Bearing Capacity and Water. Part II: Measured Response. Note 249. Road Directorate Denmark, Danish Road Institute, (3) Verstraeten, J. Bituminous Materials with a High Resistance to Flow Rutting. PIARC Technical Committee on Flexible Roads TC8,

28 Validation of Pavement Response and Performance Models Per Ullidtz Wei Zhang Susanne Baltzer Technical University of Denmark, Institute of Planning Technical University of Denmark, Institute of Planning Road Directorate, Danish Road Institute Introduction Three different types of theoretical models may be used for the prediction of pavement response and performance: 1 Analytical models: The material parameters used with these type of models are the elastic parameters (Young s modulus and Poisson s ratio) Some models include cross-anisotropic materials and/or various interface conditions (from full friction to full slip), and other models may include visco-elastic characteristics (often in the form of a Burgers model) Most models are pure response models, and rely on empirical relations between the critical response and the rate of deterioration, sometimes using an incremental-recursive procedure for predicting the gradual deterioration in terms of fatigue cracking and permanent deformation 2 Finite Element Methods (FEM): FEM may be used as an improved response model, where non-linear stress-strain relationships, or dynamic loading, may be included They may also be used to directly calculate permanent deformation based on viscous or plastic material parameters FEM can also be used with fracture mechanics or continuum damage mechanics to predict a gradual structural deterioration (fatigue) 3 Distinct Element Methods (DEM): In DEM particles are treated individually, so that the input parameters are not in the form of constitutive equations, but consist of the size and shape of the particles, the degree of compaction, the particle stiffness and the coefficient of friction between the particles With DEM there are no stresses or strains in the materials, but forces and displacements of the particles Response, permanent deformation and failure are treated in the same calculation process The response part of the analytical methods and FEM are based on solid mechanics (continuum mechanics) and may be verified through a comparison of the predicted response to a measured response 28 28

29 Measurement of Pavement Response The two main problems with respect to measuring pavement response are: the reliability of the measured values, and the durability of the measurement device When a gauge is installed in a pavement material, it will influence the stress and strain conditions in the material For a gauge and installation to be reliable, the output should be the response that would have occurred at the position of the gauge, if the gauge had not been present Durability may be a problem both during installation of the gauge and during the experiments, if they extend over some period of time Displacements During the experiments described in this paper, deflections were only measured under FWD loading at the surface These measurements are unintrusive and very accurate During previous experiments, deflections within the pavement structure has been measured using geophones or accelerometers buried in the structure To determine deflections, the signals from the geophones must be integrated once and those from the accelerometers twice Under rolling wheel loading, with a high content of low frequencies, this may be a problem An alternative is to use Multi Depth Deflectometer (MDD, de Beer et al 1988) based on Linear Variable Differential Transformers (LVDT) or LVDTs anchored at different depths (Jansson & Wiman, 1994) In either case there may be problems with both accuracy and durability Strain in Asphalt The horizontal tensile strain at the bottom of the asphalt layer is an important design parameter In 1982 the OECD initiated a Concerted Action on the Use of Large-Scale Research Facilities (OECD, 1985) As a part of this initiative, a number of different gauges for measuring strain in asphalt layers, were compared in a full-scale international experiment in Nardo in Southern Italy Based on these experiments a new asphalt strain gauge was developed at the Technical University of Denmark (OECD, 1985b) This gauge contains a number of layers to protect against heat and mechanical impact during installation and against moisture during operation This gauge normally has a 100% survival rate during installation, and some gauges are still functioning after more than seven years in an in-situ test section Stress in Soils and Unbound Materials A number of theoretical studies (Torry & Sparrow, 1966, Askegaard, 1979, Ullidtz & Busch, 1979) have shown that a pressure cell should have a high stiffness in relation to the media in which it is supposed to measure the stress (low flexibility), and that the ratio of thickness to diameter should be low (low aspect ratio) A cross section of the pressure cells, used in the experiments reported in this paper, is shown in Figure

30 Figure 1. Cross section of pressure cell. The liquid filled cavity above the gauge membrane ensures a very stiff cell The cell has been calibrated in sand (Askegaard, 1996) and in clay (Zhang, Ebersohn & Selig, 1995) under highly varying stress conditions and load histories The stress measured by the cell, q, was found to be (Ullidtz, Askegaard & Sjølin, 1996): T = 1.10 σ ( σ 2 + σ 3) where s 1, s 2 and s 3 are the major principal stresses The cells are made of titanium and all cavities are filled with silicon rubber, to ensure a high durability Six cells installed in the subgrade of three in situ test sections in Southern Sweden in 1991 were all functioning at the latest test in 1998 An in situ check on the gauge calibration can be obtained by integrating the stress bulb over the plane of the gauge, and comparing the result to the imposed load Strain in Soils and Unbound Materials Most strain measurements, both dynamic and static, were done using a gauge based on an LVDT In one test the emu system from University of Nottingham was used in parallel The agreement between the LVDT and emu strains was reasonably good, but due to a large noise to signal ratio with the emu gauges, the LVDTs were used in subsequent experiments Stresses in semi-infinite half-space Boussinesq s equations are developed for calculating stresses, strains and deflections in a homogenous, isotropic, linear elastic semi-infinite half-space An evaluation of how well the theory predicts the actual stresses and strains in a structure should therefore be carried out in a media as close to the assumptions as possible At a test in the Danish Road Testing Machine, the vertical normal stress under an FWD load was measured in a sand, with more than 90% of the grains falling within the sand fraction (60 mm to 2 mm) (Ullidtz, Askegaard & Sjølin, 1996) The stress was measured with three different gauges The sand was uniform to a depth of 700 mm and at a distance of 600 mm to either side of the gauges All gauges were installed at a depth of 30 30

31 Figure 2. LVDT based soil strain cell. 280 mm. An FWD with a 300 mm diameter loading plate was used to exert a known vertical force at the surface of the sand, with a contact stress of about 300 kpa (the exact stress was measured for each load). All the loads were imposed in the centerline of the gauges, but at different horizontal distances from the FWD to the gauges. The agreement between the three gauges was quite good, as seen in Figure 3. Figure 3 also shows the theoretical stress at the depth of the gauges. At the centerline of the loading plate the theoretical stress is only about half of the measured stress. In calculating the theoretical stress, a parabolic stress distribution was assumed. For a rigid plate on an elastic medium, the theoretical stress at the edge of the plate should be infinitely large. On a granular material this is, of course, not possible. At the edge of the plate the bearing capacity is very small, resulting in stresses close to zero. Had the theoretical stress distribution for a rigid plate on an elastic medium been used, the difference between measured and calculated values would have been even larger. The same would be the case if a uniformly distributed load was assumed. One explanation of the difference could be that the measured stresses are too large. Integrating the stress bulbs over the plane of the gauges shows that the total force measured by the gauges is 10-14% larger than the FWD peak load. This difference cannot explain the difference between measured and calculated values. Several factors that could have influenced the outcome of the measurements have been investigated: 31 31

32 Vertical Stress kpa Gauge 1 Gauge 2 Gauge 3 Bous Distance mm Figure 3. Measured and theoretical vertical stress in half-space. The sand material was not an infinite half-space but limited by a different material, both at the sides and at the bottom. Calculations with different assumptions for the modulus of the boundary material, indicated that this could not have influenced the measured values by more than a few percent. The load from the FWD is not static but dynamic. A simple uniaxial model of masses and springs was used to investigate this effect. It was found that the dynamic effects could cause the peak load at the plane of the gauges to be about 6% higher than the imposed load. This difference is in reasonably good agreement with the integration of the stress bulbs. The repeated loading with the FWD may have compacted the material in the line of the gauges to a higher modulus than the surrounding material. An FE-program showed that a 50% higher modulus in the column of material below the loading plate could increase the stress measured at the centerline by about 15%. Sand is not a linear elastic material. FEM calculations with the sand modulus being proportional to the bulk stress raised to a power of 0.5 and 1.0, showed that this non-linearity had only a minor effect on the stress distribution. The conclusion is that the measured stresses are correct, and that the theory of elasticity cannot predict the stresses with a reasonable degree of accuracy. A good agreement with measured stresses can be obtained using Harr s probabilistic stress distribution (Harr, 1977) with a coefficient of lateral stress of 0.16, or by introducing a false depth of 180 mm, equal to 0.64 times the actual depth, in the theory of elasticity

33 Stresses and Strains in Layered Structures Stresses and strains have been measured in a number of pavements in the Danish Road Testing Machine (RTM) and in situ, both under FWD loading and under rolling wheels. When trying to verify response models the following procedure has often been followed: The stresses and strains are measured at different distances from an FWD with a hydraulic pad under the loading plate (to ensure a uniform load distribution) and/or under rolling wheel loads. Preferably the whole influence line is recorded, not just the peak values. Ordinary FWD tests are done on the pavement, and the deflection basin is recorded. The layer moduli are calculated using different theoretical response models, and different assumptions about the pavement (number of layers, non-linearities, rigid bottom etc.) Using these layer moduli, possibly adjusted for differences in temperature or loading time, the response is calculated at the position of the gauges, using the same response models employed in the backcalculation of layer moduli. Measured and calculated response is compared. For a number of these verifications the program FeBack has been used. This program has three different response models: an axial symmetric Finite Element program (FE, Duncan et al. 1968, Ullidtz, 1998), a linear elastic analytical model for a layered system developed by F. van Cauwelaert for Waterways Experiment Station (WES) and the Method of Equivalent Thickness (MET) including a non-linear subgrade. An example of the input screen is shown in Figure 4. In the example below, FeBack was used with the second test pavement from the International Pavement Subgrade Performance Study in the Danish Road Testing Machine (RTM) (Ertman Larsen & Ullidtz, 1997). The main aim of this study is to develop design criteria or deterioration models for subgrade materials, but a secondary aim is to verify existing response models. For this purpose the pavement is heavily instrumented, with eight asphalt strain gauges at the bottom of the asphalt, two vertical strain cells at the bottom of the base course, and strain cells and pressure cells for measuring the vertical, longitudinal and transverse strains and stresses in the top three layers (of approximately 150 mm each) of the subgrade. Some of these measurements are also used in the AMADEUS program (Advanced Models for Analytical Design of EUropean pavement Structures). A cross section of the pavement is shown in Figure

34 Figure 4. Input screen for FeBack. With MET and FEM the pavement was treated as a three-layer system (the filter gravel and concrete bottom was disregarded). With WES the modulus of the concrete bottom was fixed at MPa, and the modulus of the native soil below the concrete was allowed to vary. Figure 5. Cross section of pavement in the RTM2 study

35 An example of layer moduli for a single test point, in MPa, is shown below. MET WES FEM E E E n E E All methods result in approximately the same asphalt modulus (E1, 25 C). The nonlinear subgrade modulus (with MET and FEM) is expressed on the form: E m σ1 = E3 p n where σ 1 is the major principal stress and p is a reference stress (here atmospheric pressure). Even though a rigid bottom is considered with the linear elastic approach (WES) the modulus of the base course (crushed gravel) is seen to be less than that of the subgrade (moraine clay). This is a typical result with linear elastic methods. For the two methods where the subgrade is considered to be non-linear, the modulus of the base course layer is much more reasonable. The peak value of the horizontal strain at the bottom of the asphalt layer, the vertical stress at the top of the subgrade, and the vertical strain at the top of the subgrade are shown in the table below. The values were measured under FWD loading, and the average measured values are given, as well as the values calculated with the three response models, for the same load. Strains are in µ/m. MET WES FE Measured Horizontal strain in asphalt, µe Vertical stress on subgrade, kpa Vertical strain in subgrade, µe All three methods are seen to predict the horizontal strain at the bottom of the asphalt layer reasonably well. This appears to be a general tendency. The vertical stress on the subgrade is overestimated by the linear elastic method, but underestimated for the two non-linear subgrades. The vertical strain in the subgrade, which is an important design parameter, is underestimated by a factor of two by the linear elastic method. The figure below shows the influence lines for vertical strain at the top of the subgrade, as measured by four different gauges and as predicted by the three different response models. In this case MET is seen to result in the best prediction. This has been found to be the case also in many other instances

36 Vertical strain at top of subgrade Distance from load centre, m Figure 6. Measured and calculated vertical strain in subgrade. MET WES FE None of the methods could correctly predict the vertical strain at the bottom of the base course. Pavement Performance Prediction of pavement performance from pavement response is the core of the analytical-empirical method of pavement design. Pavement performance may be composed by different measures of the functional condition (ride quality, safety) and structural condition (bearing capacity). This paper discusses models to estimate the plastic strain at the top of the subgrade and permanent deformation in the subgrade. Models for estimation of pavement deterioration are also studied. Models for Estimation of Plastic Strains at the Top of the Subgrade A maximum permissible normal strain or stress at the top of the subgrade is used in pavement design based in the analytical-empirical method. Most of the empirical relations between load repetitions and permissible stress or strain being used today are of type: permissibl e stress or N strain = K * a E * E' b (5.1) where N is the number of load repetitions to cause a certain deterioration, E is the modulus of a material, E is a reference modulus, and K, a and b are constants or functions of the materials and environmental condition

37 The permissible stress or strain in the equation above is normally obtained either from laboratory tests or from in-situ observation, and is based on the normal stress or strain (resilient value). For several test pavements built in the Danish Road Testing Machine, the resilient and plastic strains in the subgrade layers were recorded after different number of load repetitions with LVDTs. This makes it possible to establish a model for the development of plastic strain in the subgrade. Given a permissible permanent deformation of the subgrade, the corresponding number of load repetition may then be calculated. The assumption is made that the vertical plastic strain, e pz, can be written as a function of the number of load repetitions, N; and of the resilient response in the following format: the strain model the stress model the energy density model ε ε ε pz pz pz = A * N = A* N = A * N α α α * ε β z 1 σ z * 2 p β σ z * p * ε β z (5.2) where ε pz is the vertical plastic strain in microstrain at depth z, ε z is the vertical resilient strain in microstrain at depth z, σ z is the vertical stress at depth z, p is a reference stress (here atmospheric pressure), and A, a and b are constants. The increment of the plastic strain at different load levels may be calculated using an incremental-recursive procedure based on: the strain model the stress model the energy density model dε dε dε pz pz pz = α* ε = α* ε = α* ε p p p A* ε ε pz β z σ z A* p ε pz 1 α * dn 1 α 1 σ z A* z ε 2 p ε pz β β * dn 1 α * dn 37 37

38 Two test pavements for the International Pavement Subgrade Performance Study in the Danish Road Testing Machine (RTM) were constructed, instrumented and loaded. Based on the measured plastic strain in the subgrade layers of the first test pavement (RTM1), constants A, α and β for the strain-, stress-, and energy density model were computed using a minimization process in which the difference between the measured and estimated plastic strain were minimized. Then these constants were used to predict the plastic strain in the subgrade layers of the second test pavement (RTM2). The constants A, α and β for the strain-, stress-, and energy density model are given below. A α β Strain-model Stress-model Energy density -model Figures 7 and 8 show the measured and estimated vertical plastic strain in the first subgrade layer of RTM1 and RTM2, respectively Microstrain Measured strain stress energy Load repetition Figure 7. Measured and estimated vertical plastic strain in the first subgrade layer of RTM1. Microstrain Load repetition Measured strain stress energy 38 Figure 8. Measured and estimated vertical plastic strain in the first subgrade layer of RTM2. 38

39 The stress-, strain- and energy density models are also used to estimate the plastic strain in subgrade layers 2 and 3, and the estimated values are compared to the measured values. The result shows that the energy density model produces a reasonable estimation of plastic strain in the first subgrade layer for both RTM1 and RTM2. Model for Estimation of Permanent Deformation in the Subgrade One of the objectives of the International Subgrade Performance study is to establish a subgrade rutting model. The permanent deformation in the subgrade is related to the plastic strain, so the energy density model is used to compute permanent deformation in the subgrade. Equation (5.2) can also have the following format, if the vertical resilient strain at depth z (ε z ) is substituted by the ratio of the vertical stress (σ z ) and the resilient modulus (E). ε pz = A* N α 2 σ * z 2 E p β *10 6( β 1) (5.3) where strains are in m/m (in equation (5.2) strains are given in µm/m). The relationship between the plastic strain and plastic modulus can be expressed as follow: E E pz σ = ε substituti ng pz = A* N = A* N = C α α ε σ z * p p z pz pz σ z β 2 σ z 2 E p with equation (5.3) then gives *10 β (2C) ( β 1) * p *10 n p 6( β 1) 6( β 1) σ ( nβ + 1 2β) z where where σ z E = C p n p n = nβ + 1 2β (5.4) where p C p n p is a reference stress (atmospheric pressure), and is a material constant at a given number of load repetitions, is a material constant. Equation (5.4) gives a relationship between the plastic modulus and vertical normal stress. The relationship between the permanent deformation (d pz ) and vertical plastic strain (ε z ) in the subgrade will have the following format: This formula is derived from Boussinesq s vertical displacement equation, when the constant modulus E is substituted by a non-linear function of the major principal stress

40 d pz ε pz = (1 2 * n p * z ) (5.5) From equation (5.3) the plastic strain at different depth in the subgrade can be expressed in the following format: ε ε pz1 pz2 = A* N A* N α α σ = σ 2 σ z1 * 2E p 2 σ z 2 * 2E p z1 z 2 2β β β z = z *10 * ( β 1) 6( β 1) 4β ε pz1 z = z 2 1 4β * ε p z2 (5.6) Using the preceding equations, a plastic strain at the top of the subgrade can be estimated when a permissible permanent deformation in the subgrade is given. Applying the energy density model (equation 5.2) and a computed resilient strain and stress at a certain load level, the permissible number of load repetitions can then be calculated. The two following cases are examples of this type of computation. Case 1: Input: permissible permanent deformation at the top of the subgrade is 10 mm; resilient strain is 8000 microstrain, (a 60 kn dual wheel load on RTM1); resilient stress is 160 kpa, (a 60 kn dual wheel load on RTM1); equivalent thickness to the subgrade 600 mm; elastic non-linearity parameter, n=-0.47 plastic non-linearity parameter, n p =nβ+1-2β=-0.47* *0.868=-1.144; parameters in the model to estimate the plastic strains, A=0.453, α=0.341 and β=0.868 (the energy density-model); Output: the plastic strain at the top of the subgrade (using equation (5.5)) d pz(1 2n p) 10 mm * (1 2 *( 1.144)) ε = = = z 600mm pz µ strain the permissible number of load repetitions (using equation (5.2)) 40 40

41 N α ε = 1 A* * ε 2 pz ε N = 1 A* * ε 2 pz z z σ z σ ' σ z ' σ β β 1 α = *8000 * = 164,729 To produce 10 mm rutting at the top of the subgrade in RTM1, a 60 kn wheel load with approximately 165,000 load repetitions is needed. This is a reasonable prediction considering the following facts for RTM1 after 100,000 load repetitions (a 20 kn wheel load with 50,000 load repetitions and 40 kn wheel load with 50,000 load repetitions): the permanent deformation at the top of the subgrade was approximately 4 mm; the resilient strain at the top of the subgrade was about 5300 microstrain. Case 2: Input: permissible permanent deformation at the top of the subgrade is 10 mm; resilient strain is 4500 microstrain, (a 60 kn dual wheel load on RTM2); resilient stress is 160 kpa, (a 60 kn dual wheel load on RTM2); equivalent thickness to the subgrade 570 mm; elastic non-linearity parameter, n=-0.57 plastic non-linearity parameter, n p =nβ+1-2β=-0.57* *0.868=-1.231; parameters in the model to estimate the plastic strain, A=0.453, α=0.341 and β=0.868 (the energy density-model); Output: the plastic strain at the top of the subgrade (using equation (5.5)) d pz(1 2n p) 10mm*(1 2*( 1.231)) ε = = = z 570mm pz µ strain the permissible number of load repetitions (using equation (5.2)) N α ε pz = 1 A* * ε 2 ε pz N = 1 A* * ε 2 z z σ z σ ' σ z ' σ β β 1 α = * 4500 * = 963,

42 To produce 10 mm rutting at the top of the subgrade in RTM2, a 60 kn wheel load with approximately 960,000 load repetitions is needed. This is a reasonable prediction based on the following facts for RTM2 after 150,000 load repetitions (a 40, 50 and 60 kn dual wheel load with 50,000 load repetitions each): the permanent deformation at the top of the subgrade was approximately 3 mm; the resilient strain at the top of the subgrade was about 4500 microstrain. Table 1 summarizes the measured and estimated values of RTM1 and RTM2 used in the estimation for the number of load repetitions, which will produce a permanent deformation of 10 mm at the top of the subgrade. RTM1 RTM2 Dual wheel load Load repetitions (kn) (thousand) Total load repetitions 100, ,000 Measured vertical resilient strain in SGL1 (µe) Plastic strains at the top of the subgrade (µe) Permanent deformation at the top of subgrade (mm) Vertical resilient strain used in the estimation (µe) Vertical stress used in the estimation (kpa) Load repetitions to produce 10 mm permanent deformation at the top of the subgrade, at 60 kn 165, , measured under the 40 kn dual wheel after 100,000 load repetitions. measured under the 60 kn dual wheel after 150,000 load repetitions. computed from the measured strain in the first subgrade layer under a 60 kn dual wheel. Table 1 The measured and estimated values of RTM1 and RTM2 used in the estimation for load repetitions The number of load repetitions to produce 10 mm permanent deformation at the top of the subgrade for RTM1 is much less than for RTM2. The large difference in the number of load repetitions is mainly due to the vertical resilient strains in the subgrade. Relatively large vertical resilient strains in subgrade layer 1 of RTM1 were recorded. The vertical resilient strain at the top of the subgrade used in the estimation is 8000 microstrain for RTM1 and 4500 microstrain for RTM2. The deterioration rate of RTM1 is much higher than that of RTM2. Models for Predicting Roughness and Rut Depth of Pavements Longitudinal and transverse surface profiles of RTM1 and RTM2 were observed at different number of load repetitions. Roughness measured with Slope Variance (SV) and International Roughness Index (IRI) was obtained from the measured longitudinal profiles, and Rut Depth (RD) was computed from the measured transverse profiles

43 A model to predict SV, IRI and RD was established based on the measured vertical resilient strain and stress in the subgrade. The assumption made in the models was that SV, IRI and RD are functions of the number of applied load repetitions and of the internal energy density. The model has the following format: SV ( IRI or RD) = A* N α 1 * ε 2 z σ z p β the energy density model The increment of SV (IRI or RD) at different load levels may then be calculated using an incremental-recursive procedure based on: 1 σ z A* z ε 2 p d ( SV ) = α * SV * SV β 1 α * dn the energy density model where SV may be substituted by IRI or RD, ε z is the vertical resilient strain at the top of the subgrade, σ z is the vertical stress at the top of the subgrade, p is a reference stress (atmospheric pressure, 0.1 MPa), and A, α and β are parameters derived from the experimental data. Using this procedure, parameters A, a and b for SV, IRI and RD are obtained based on the experimental data collected on RTM1, and listed in the following table. Parameters A, α and β for SV, IRI and RD. A α β Energy-model SV 1.92* IRI 5.92* RD 3.03* Figure 9 to 11 shows the measured and predicted SV, IRI and RD of RTM1, respectively, while figure 12 to 14 shows the measured and predicted SV, IRI and RD of RTM2, respectively

44 Measured predicted SV Load repetition Figure 9. Measured and predicted SV for RTM IRI m/km Measured predicted Load repetition Figure 10. Measured and predicted IRI for RTM1. 12 Rut depth mm Measured predicted Load repetition Figure 11. Measured and predicted RD for RTM

45 10 8 Measured predicted 6 SV Load repetitions Figure 12. Measured and predicted SV for RTM2. 5 IRI m/km Measured predicted Load repetitions Figure 13. Measured and predicted IRI for RTM2. Rut depth mm Measured predicted Load repetitions Figure 14. Measured and predicted RD for RTM

46 With the energy density model the predicted SV, IRI and RD fit the measured values reasonably well. Conclusion Currently used models for calculating pavement response are based on a number of simplifications. Although some of the more advanced Finite Element models may include dynamic loading, viscous and plastic materials characteristics and the dependency of material characteristics on stress or strain conditions, they are still based on continuum mechanics, whereas most pavement materials have discontinuities and many even are particulate. The validity of these models must, therefore, be experimentally verified. Experiments, where theoretical stresses and strains have been compared to measured values, show some of the limits of the existing methods. But there are also indications that some of the important design parameters (horizontal strain at the bottom of an asphalt layer, vertical strain at the top of the subgrade) can be predicted with a reasonable degree of accuracy, even when using highly simplified methods like the Method of Equivalent Thicknesses (MET). There appears to be a need for developments on two fronts: One is the development of more realistic models, with a reduced discrepancy between the nature of real pavement materials and the assumptions on which the models are based. These sophisticated models are needed to get a better understanding of the behavior of pavement materials and pavement structures. The other need is for models that may be used for every day evaluation of pavement structures. These simplified models are needed for the practical work of highway engineers. On both fronts, however, it is of paramount importance that the models be verified against the reality of actual pavement materials and actual road structures. References Askegaard, V. Pressure Cell Calibration Tests, Report No 9508, Series S, Department of Structural Engineering and Materials, The Technical University of Denmark, 1996 Askegaard, V. Stress and strain measurements in solid materials, Report No. R 92, Dept. of Structural Engineering, Technical University of Denmark, de Beer, M, Horak, E. & Visser, A.T. The Multi-depth Deflectometer (MDD) system for determining the effective elastic moduli of pavement layers, First International Symposium on nondestructive testing of pavements and backcalculation of moduli, Baltimore, June

47 Duncan, J.M., Monismith, C.L. & Wilson, E.L. Finite Element Analysis of Pavements, Highway Research Record 228, Ertman Larsen, H.J. & Ullidtz, P. Pavement Subgrade Performance Study in the Danish Road Testing Machine, Eighth International Conference on Asphalt Pavements, Proceedings Vol. 1, University of Washington, Seattle Harr, M.T. Mechanics of Particulate Media, McGraw-Hill, New York, Jansson, H. & Wiman, L.G. Mätta och beräknade deformationer i vägen vid fallviktsmätning ( Pavement analysis based on measured in-depth deflection data ), VTI meddelande Nr 738, 1994 (in Swedish, summary in English). OECD Full-scale Pavement Tests, Road Transport Research, Organisation for Economic Co-operation and Development, Paris, 1985a. OECD Strain Measurements in Bituminous Layers, Organisation for Economic Cooperation and Development, Paris, 1985b. Tory, C. & Sparrow, R.W. Behaviour of a Soil Mass under Dynamic Loading, Journal of the Soil Mechanics and Foundation Division, Proceedings of the American Society of Civil Engineers, Ullidtz, P. & Busch, C. Laboratory Testing of a Full Scale Pavement: The Danish Road Testing Machine, Transportation Research Record 715, Transportation Research Board, Ullidtz, P., Askegaard, V. & Sjølin, F.O. Normal Stresses in a Granular Material Under Falling Weight Deflectometer Loading, Transportation Research Record 1540, Transportation Research Board, Ullidtz, P. Modelling Flexible Pavement Response and Performance, Polyteknisk Forlag, Lyngby Zhang, J., Ebersohn, W & Selig, E.T. Evaluation of Earth Pressure Cells For Pavement Subgrade Performance Study, Geotechnical Report No. CRR95-427F, University of Massachusetts,

48 Modeling Pavement Response and Estimating Pavement Performance Wei Zhang Robin A. Macdonald Technical University of Denmark, Institute of Planning Road Directorate, Danish Road Institute Abstract Two test pavements in the Danish Road Testing Machine (RTM) have been successfully constructed, instrumented and load tested for the International Pavement Subgrade Performance Study. One of the objectives of the study is to establish a subgrade failure criterion. The subgrade failure is related to the subgrade mechanical properties, loading conditions and moisture contents of the subgrade soils. Two test pavements instrumented to measure soil suctions, temperatures and responses of the pavement materials were constructed. Pavement response and pavement surface profiles were recorded during the accelerated loading test. This paper describes the test pavement design, the instruments used to measure pavement responses and the installation of instruments. Some problems encountered with back-calculation are discussed. The measured and calculated pavement responses are compared. Based on measured plastic strains in the subgrade layers and measured surface profiles, models for estimating plastic strains at the top of the subgrade, permanent deformation of the subgrade and pavement deterioration have been established. Introduction Currently, flexible pavements are designed to withstand a maximum horizontal tension at the bottom of the asphalt layer and a maximum vertical compressive strain (or stress) at the top of the subgrade. These maximum strain values are chosen to limit fatigue damage of the asphalt layer and rutting (permanent deformation) of the subgrade. The subgrade is assumed to be the layer in which the predominant amount of rutting occurs. The criterion for allowable subgrade strain was originally developed by Dormon and Metcalf [ref.1] based on data from the American Association of State Highway Officials (AASHO) test road. As the AASHO test loops were built on the native soil (silty clay, A-6) in Ottawa, Illinois, USA, no data were available to describe the behavior of different types of soils. Also, the failure criterion does not distinguish between wet and dry soil conditions which show very different strengths. In most cases, the only input to the current subgrade failure criterion is traffic. The existing failure criteria has the following format: 48 48

49 σ a ( or ε a N ) = K * a E * E' b (1.1) where σ a is the allowable stress, ε a is the allowable strain (horizontal strain at the bottom of the asphalt layer for cracking or vertical strain at the top of subgrade for rutting.), N is the number of loads to cause a certain deterioration. E is the resilient modulus of the material, E is a reference modulus, and K, a and b are constants, or functions of materials and environmental conditions, and obtained from either laboratory tests or long term observations on a road network. Material type and degree of saturation do have an influence on the strength (resilient behavior, stiffness) and thereby resistance to rutting (permanent deformation under repeated loading) of soils. The study summarized in this paper is part of an international project, the Pavement Subgrade Performance Study, which sets out to investigate how rutting develops for different conditions and how this can be used in pavement design. In the main part of the Subgrade Study, which is being conducted in the USA, full-scale test sections are being constructed with four different subgrade materials. These pavements are instrumented so that the stresses, strains, temperatures and moisture contents at several positions can be measured as accurately as possible. The pavement sections are being trafficked by an automated loading facility (an Heavy Vehicle Simulator, HVS, developed and manufactured in South Africa) and the subsequent rutting will be monitored closely and recorded. Also as part of the project, laboratory testing will evaluate the resilient and plastic properties of the subgrade materials used in the test pavements. When the accelerated loading testing has been finalized, the rutting criterion developed will be validated on in-service test sections in Finland (VTT) and in the U.S.A. (the Mn/ROAD Project). The broad objectives of the Pavement Subgrade Performance Study described in the Workplan [ref.2] are: (a) (b) (c) Develop an improved mechanistic subgrade failure criterion (elastic or plastic) for new and reconstructed pavements, and integrate the study findings into improved mechanistic-empirical design methodologies for new and reconstructed flexible pavements. Develop a solid basis for future research on seasonal effects of environment on pavements. Evaluate the effect of environment on resilient material properties, especially the effect of moisture content over time in base and subgrade layers, i.e., the seasonal variability of pavement materials

50 (d) As a corollary objective, to foster international co-operation in pavement research between agencies, including the U.S. Federal Highway Administration (FHWA), the U.S. Army Corps of Engineers (CRREL), the Minnesota Department of Transportation (Mn/ROAD), the Cornell University Local Roads Program (CLRP), the Danish Road Institute (DRI) (of the Road Directorate, Ministry of Transport of Denmark), and the Technical Research Centre of Finland (Finnish National Road Administration of the Ministry of Communication, Transport, and Nordic Co-operation of the Republic of Finland). Two test pavements for the International Pavement Subgrade Performance Study have been constructed, instrumented and load tested in the Danish Road Testing Machine (RTM). This paper describes the experiment design, the instrumentation, the measured and calculated pavement responses, and the measured and estimated pavement performances of the RTM1 and RTM2 test pavements. Design of Experiment Danish Road Testing Machine The Danish Road Testing Machine (RTM) is a pavement testing facility, with a width of 2.5 m (8.2 ft.) and a length of 27 m (88.6 ft.). The central 9 m (29.5 ft.) is the actual test section, which is 2 m (6.56 ft.) deep. Plan and section schematics of the RTM are shown in Figure 1. Figure 1. The Danish Linear Track Road Testing Machine (1.0 m = 3 ft in.) 50 50

51 The RTM is enclosed in a climate chamber, 4 m (13.1 ft.) wide and 3.8 m (12.5 ft.) in height. Heating and cooling equipment make it possible to maintain a temperature range of -10 C (14 F) to + 40 C (104 F). The ground water level is automatically controlled and may be raised or lowered as desired. The wheel load is hydraulically applied by a single or a dual wheel. The maximum dual wheel load is 65 kn (14.6 kip) and the maximum velocity is about 25 km/h (15.5 mph). Approximately 5,000 load repetitions may be applied during a normal working day at this load level. Since the 65 kn (14.6 kip) dual wheel load is a halfaxle load, using a fourth power relationship, this corresponds to approximately 35,000 passes of a standard 80 kn (18 kip) axleload. The lateral position of the wheel can be automatically changed during testing to give a desired transverse wheel load distribution (wander). In the current project a dual wheel load only is being applied. The speed of the dual wheel load is approximately 20 km/h (12.4 mph). The speed of transverse movement (wander) of the dual wheel load is set at approximately 0.2% of the driving speed. The width of the transverse movement (wander) is about 0.90 m (35.4 in.). Test Pavement Design and Construction The two test pavements (RTM1 and RTM2) were carefully designed to produce a certain amount of rutting within a relatively short timeframe (two months). The design, documented in the report Analysis and Prediction of Pavement Response in the Danish Road Test Machine (ref. 3), is based on the following considerations: The applied dual wheel load was planned to be at two levels: 35 kn (7.87 kip) with 100,000 load repetitions and 50 kn (11.24 kip) with 25,000 load repetitions. It was estimated that after applying these load levels and corresponding number of load repetitions, a rut depth of approximately 25 mm (1 in.) would be developed at the surface of the test pavement. The pavement structure was set at 75 mm (3 in.) for the asphalt concrete, 150 mm (6 in.) for the granular base course, 1300 mm (51.2 in.) for the subgrade and 200 mm (8 in.) for the granular drainage layer. The subgrade modulus was assumed to have a value of 50 MPa (7.25 ksi), based on the laboratory test data, and the modular ratio of the base course modulus to the subgrade modulus initially adopted was approximately 3. The asphalt modulus was assumed to have a value of approximately 3000 MPa (435.1 ksi). It was intended to generate rutting in the RTM test pavement without generating significant cracking or roughness in the asphalt concrete surface. With the assumed pavement structure, the assumed elastic moduli of the pavement layers and the planned load levels, the expected vertical strain at the top of the subgrade, and the expected horizontal strain at the bottom of the asphalt were calculated, using multilayer linear elastic theory. The expected vertical strain at the 51 51

52 top of the subgrade was initially determined to be about 1100 microstrain with a dual wheel load of 35 kn (7.87 kip), and about 1500 microstrain with a dual wheel load 50 kn (11.24 kip). The expected horizontal strain at the bottom of the asphalt would be in the range of microstrain with a dual wheel load of 35 kn (7.87 kip), and approximately 500 microstrain with a dual wheel load 50 kn (11.24 kip). The existing surface and subgrade failure criteria were utilized to obtain the permissible vertical strain at the top of the subgrade and the permissible horizontal strain at the bottom of the asphalt. After construction, FWD tests were carried out on the asphalt surfaces of the two test sections. Using the actual layer thicknesses and the layer moduli back-calculated from the FWD data, the load levels eventually applied were adjusted to produce the desired vertical strain at the top of the subgrade. A dense-graded asphalt concrete is used for the asphalt concrete layer. The maximum aggregate size is 16 mm (0.63 in). The AC mix is composed of crushed granite aggregate, sand and fillers with a 60 pen. bitumen binder. Mix tests recommended a bitumen content of 5.0 % and a Marshall density of 2420 kg/m 3 ( lb/ft 3 ) for RTM1, and a bitumen content of 5.3 % and a Marshall density of 2420 kg/m 3 ( lb/ft 3 ) for RTM2. The base course is a well-graded crushed natural aggregate with an aggregate particle size ranging from 0-32 mm ( in.). It has an optimum moisture content of 7.0 % and a maximum dry density of 2200 kg/m 3 ( lb/ft 3 ), as determined in a vibratory density test developed by the DRI. The level of compaction varied throughout the pavement between 94%-97 % of the maximum dry density, with moisture contents of between 4%-5%. A Danish Moraine Clay classified as a clayey silty sand (AASHTO classification A- 4(0)) is being used as the subgrade soil. It has an optimum moisture content of 9.0 % and a maximum dry density of 2045 kg/m 3 ( lb/ft 3 ) determined in the standard Proctor density test. It is constructed in nine layers of approximate 150 mm (6 in.) thickness each. The reinforced concrete bottom was tested with an FWD before construction of the filter gravel layer. Based on the previous studies carried out in the RTM, the concrete bottom has a modulus of 27,500 MPa (3,989 ksi) and the drainage layer has a modulus of 250 MPa (36.26 ksi). Response to be Measured and Instrument Selection In order to establish subgrade failure criterion, the development of stresses and strains in the subgrade along with the number of load receptions should be known; additionally, the vertical stresses and strains at the top of the subgrade are of particular interest for the design criterion. Based on the state of the practice, a maximum horizontal strain at the bottom of the asphalt layer is used for flexible pavement de

53 sign. Thus, the longitudinal and transverse strains at the bottom of the asphalt layer should also be measured. To measure the longitudinal and transverse strains at the bottom of the asphalt layer, Asphalt Strain Gages (ASGs, shown in Figure 2) manufactured in the RTM workshop were selected. Figure 2. Asphalt Strain Gage (25.4 mm = 1 in.) ASGs consist of a commercial strain gage (120 ohm) sandwiched between two flexible strips of fiber reinforced epoxy. The epoxy strip is fixed at both ends to 75 mm (3 in.) long stainless steel anchors. The gages are well anchored in the bituminous layer so that when there is tension at the bottom of the asphalt concrete, the distance between the anchors increases, and the strain gages are stretched. When the strain gage is stretched, its electrical resistance increases, and this change can be recorded and converted into strain. The initial distance between the anchors is approximately 100 mm (4 in.). Eight ASGs were installed longitudinally and transversely (four in each axis) in the RTM1 and RTM2 test pavements. Soil Pressure Cells (SPCs, shown in Figure 3) are used to measure stresses in the subgrade. The SPCs used in the RTM have been manufactured in the RTM workshop. They are manufactured in titanium and have their flat surfaces coated with epoxy and sand to aid good contact with primarily the subgrade soil to minimize slipping. The diameter of the cell is 68 mm (2.68 in.). The SPC is fabricated as a double diaphragm cell, in which two membranes are utilized. The cavity between these two membranes is filled with an incompressible oil (hydraulic liquid). A rosette strain gage is bonded to the bottom of the measuring membrane (the small membrane) and produces an electrical signal whenever the membrane deforms. The four strain gages in the rosette are coupled in a Wheatstone bridge circuit at a bridge voltage of 12 Volts DC. Further details of the pressure cell are given in reference [4]. In RTM1, SPCs were installed to measure horizontal and vertical stresses, while in RTM2, only vertical stresses were recorded