Long-term shear strength of geosynthetic clay liners

ARTICLE IN PRESS

Geotextiles and Geomembranes 26 (2008) 130–144 www.elsevier.com/locate/geotexmem

Long-term shear strength of geosynthetic clay liners Werner Mu¨ller, Ines Jakob, Stefan Seeger, Renate Tatzky-Gerth Federal Institute for Materials Research and Testing (BAM), D-12200 Berlin, Germany Received 17 May 2007; received in revised form 29 August 2007; accepted 30 August 2007 Available online 26 December 2007

Abstract Geosynthetic clay liners (GCLs) often have a sandwich-like multilayer structure, e.g. bentonite encased between two geotextile layers connected by fibers or yarns, either by needle-punching or stitch-bonding. Therefore, the internal shear strength of the GCL depends on the strength of reinforcing fiber bundles or yarns and their anchoring strength in the cover and carrier geotextiles. When used on long and steep slopes and covered with thick soil layers, the GCL is permanently exposed to a combined action of compressive and shear stress. Such load conditions are characteristic for landfill covers and the slope stability of the overall cover system in the long run strongly depends on the long-term internal shear strength of the GCL. A new test method was developed to study this long-term shear behavior. The focus was not only on creep, as it is normally done, but on aging effects. The shear test devices allow the measurement of creep curves and times-to-failure at elevated temperatures in different media (tap water and de-ionized water). In this publication, the main findings of the experiments on needle-punched GCLs with and without thermal treatment are summarized. Tap water as a test medium was essential to ensure sodium to calcium ion exchange in the bentonite layer. Under this condition extremely long test durations without failure were achieved. Sliding failure occurred when de-ionized water was used. Two failure modes were observed: brittle failure of the GCLs with thermal treatment and slow disentanglement of fiber bundles for untreated GCLs. Short-term shear strength (e.g. peel strength) is unrelated to the actual long-term shear strength, i.e. to the times-to-failure achieved in long-term shear strength test. Hence, short-term shear strength alone will not provide reliable dimensioning data for product design and choice of resins. Therefore, the often suggested approach, namely, restriction to short-term tests only and application of factors of safety, is challenged by these results. r 2007 Elsevier Ltd. All rights reserved. Keywords: Geosynthetic clay liner; Shear strength; Oxidative resistance; Test methods; Long-term testing

1. Introduction In many cases a geosynthetic clay liner (GCL) is a sandwich of a carrier and cover geotextile and a layer of wet or dry powder or granular sodium or calcium bentonite between. The geotextile may be a nonwoven, a woven or a geocomposite of both types. The system geotextile/ bentonite/geotextile is connected and reinforced by needlepunching or stitching (Koerner, 1994). The advantages of GCLs compared to various other liner materials are their easy transportation and installation, robustness against typical construction site stresses and cost effectiveness. However, the qualification as a barrier to water flow in geotechnical liner systems requires the proof of long-term low water permeability and long-term high internal shear Corresponding author. Tel.: +49 30 8104 1432; fax: +49 30 8104 1437.

E-mail address: [email protected] (W. Mu¨ller). 0266-1144/$ - see front matter r 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.geotexmem.2007.08.001

strength (Simon and Mu¨ller, 2004). An overview about properties and different applications of GCLs is given elsewhere (Bouazza, 2000; Zanzinger et al., 2002). The issue of long-term internal shear strength is of prior importance when GCLs are used to line long and steep slopes, for example in landfill covers. On slopes the GCL is permanently exposed to shear stresses caused by the downhill component of the gravitational force which acts on the drainage and reclamation layer on top of the geosynthetic. Therefore, the sliding stability of the overall cover system depends on the friction forces between the GCL product and the adjacent cover system components as well as on the internal shear strength within the GCL itself. The actual shear stress encountered in a specific application is determined by the configuration of the cover system (Dickinson and Brachmann, 2006). In rare cases the GCL product is installed over an uneven subgrade and constrained sufficiently by covering soil. Then, the

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adjacent surfaces will interlock, in which case shear and friction forces are no significant factors. In most cases— particularly in composite cover systems made up of several geosynthetic layers—clearly defined sliding planes emerge. Then, the forces acting down the incline must fully be absorbed by the internal shear resistance of the cover system components and the friction forces at the interfaces between them (Vukelic´ et al., 2008). Sodium bentonite of high water content has a very low shear resistance. Therefore, the shear strength of the GCL depends on the strength of the reinforcing fiber bundles or yarns and their anchoring strength in the cover and carrier geotextiles. For polyolefin geotextile materials, it is not only creep but also degradation by stress cracking and oxidation which might impact shear strength of the GCL. Creeping is an intrinsic property of an intact polymeric material. Under load, the polymer molecules continuously reorganize their arrangement to fit to the external constraint. The reorganization is accompanied by a decrease in entropy and the required increase in free enthalpy is delivered by the work done by the external force. Under creep cross-sections are continuously reduced and the local stress might eventually become so large that ductile fracture will occur. In the long run, the material is changed by aging processes either of physical or chemical nature. These changes strongly influence mechanical strength and creep behavior. Brittle fracture might occur well below the stress and strain limits which were derived from creep curve measurements. Environmental stress cracking (ESC) is the most relevant example of a physical aging process of polyolefin plastics. Under load, micro-cracks might be formed between structural elements constituting the morphology of the polymeric material. These micro-cracks can grow and eventually form macro-cracks which will cause brittle failure (Mu¨ller, 2007). The orientation of the fiber material strongly reduces the sensitivity to stress cracking. However, even in oriented fibers stress crack like phenomena may occur, i.e. crack formation in the long run at low stress level (Seeger et al., 2002). When the fibers are thermally treated and even locally melted the orientation is destroyed and the molten parts of the fibers become susceptible to ESC. Polypropylene (PP) and polyethylene (PE), which are often used in geotextiles, are susceptible to the chemical aging process of oxidative degradation. When exposed to oxygen, the polymer molecules can be decomposed by a radical chain reaction leading to a decrease in strength and an acceleration of creep. Antioxidants, which interrupt oxidative chain reactions, are added to the polymer resins to ensure oxidation stability. Factors, like polymer structure, residues from the polymerization process, type and amount of antioxidants, design of the geosynthetic, depletion of antioxidants by migration and oxidative consumption and the availability of oxygen under application conditions, have strong influence on the service life of PP or PE geosynthetics, which may vary by orders of

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magnitude (Mueller et al., 2003). In complex-structured geocomposites, like GCLs, failure mechanisms might occur which are neither a consequence of creep nor that of aging. Slow disentanglement of fibers, for example, should be considered as a potential failure mode. Meanwhile, results of field performance studies conducted over about 10 years (Bonaparte et al., 2002) and long-term laboratory creep tests (up to 10,000 h) on GCLs at room temperature (RT) are available (Koerner et al., 2001; Zanzinger and Alexiew, 2002). Based on models for visco-elastic behavior these creep data have been extrapolated to service lives of 100 years. Such extrapolations assume that for 100 years the above-mentioned aging processes (oxidative degradation, ESC and enhanced disentanglement) are not relevant, which in general is not true. For this reason a long-term creep test is needed which considers the combined influence of aging and creep. To accelerate aging, it is necessary to perform such a test at elevated temperatures. In this publication, results of such long-term shear strength tests on needle-punched GCLs are presented. A shear test device, which was already successfully applied to textured geomembranes (Seeger et al., 2000), was adapted to study the shear behavior of some types of non commercial grade, which were exclusively produced for test purposes, and commercially available needle-punched GCLs (Thies et al., 2002). A similar approach was suggested by Hsuan and Koerner (Hsuan, 2002). 2. Experimental 2.1. Sample properties Samples from four different types of GCLs, all reinforced by needle-punching, were tested. The nonwoven geotextiles were made of either high-density PE (PE1) or of two different PP resins (PP1 and PP2). The slit film woven geotextiles used in the carrier geocomposite were made of three different PP resins (PP3, PP4 and PP5) and of a highdensity PE (PE2) resin. Two different manufacturing methods were applied. Method A resulted in an enhancement of the anchor strength of the reinforcing fiber bundles in the carrier geocomposite compared to method B. Enhancement was achieved by thermal treatment, i.e. flame burnishing the backside of the carrier geocomposite. The samples were composed of these components and denoted as given in Table 1. The mass per unit area of the cover nonwoven geotextile was 300 g/m2. The carrier geocomposite was composed of a 250 g/m2 nonwoven geotextile and of a slit film woven geotextile of about 100 g/m2 connected by needle-punching. The mass per unit area of the dry sodium bentonite powder was about 4700 g/m2 with the exception of sample GCL 2Ac. This sample was a needle-punched GCL with a double bentonite layer separated by a nonwoven geotextile. The mass per unit area of both bentonite layers was about 6800 g/m2. The carrier geotextile of sample GCL 2Ad was

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Table 1 Sample designation and manufacturing method Sample designation

GTnw resin

GTw resin

Manufacturing method

Description

Peel strength (N/(10 cm))

GCL 1A GCL 1B GCL 2Aa GCL 2Ab GCL 2Ac GCL 2Ad GCL 2B

PE1 PE1 PP1 PP2 PP1 PP2 PP1

PE2 PE2 PP3 PP4 PP3 PP5 PP3

A: thermal fixed B: not thermal fixed A: thermal fixed A: thermal fixed A: thermal fixed A: thermal fixed B: not thermal fixed

GTnw/Na-bentonite/GC GTnw/Na-bentonite/GC GTnw/Na-bentonite/GC GTnw/Na-bentonite/GC GTnw/Na-bent./GTnw/Na-bent./GC GTnw/Na-bentonite/GTw GTnw/Na-bentonite/GC

230730 214727 119719 163743 111714 110717 60727

Italic: products commercially not available, GTnw: nonwoven geotextile, GTw: woven geotextile; GC: geocomposite.

only a 100 g/m2 woven geotextile. The samples differed in their short-term peel strength due to differences in materials and manufacturing methods (Table 1). The peel strength was measured by adapting the wide width tensile test DIN EN ISO 10319 (von Maubeuge and Ehrenberg, 2000; von Maubeuge and Lucas, 2002). Short-term internal shear strength was determined by shear box tests following EN ISO 12957-1. A value of about 5576 kPa (after 24 h hydration at confining normal test stress of 30 kPa) was found for specimens from a sample equivalent to GCL 2Aa and of about 4872 kPa (after 48 h hydration at confining normal test stress of 20 kPa) for sample GCL 2Ad. As will be discussed later, the specimens were loaded in the shear stress test devices with a shear stress of 17 kPa and a normal stress of 43 kPa up to a temperature of 80 1C in a water bath. In none of the experiments, there was any indication of immediate sliding. Therefore, the short-term internal shear strength of the GCL specimens was well above 17 kPa at normal stress of 43 kPa at 80 1C.

Fig. 1. Long-term shear strength test device (Seeger et al., 2000).

2.2. Long-term shear strength testing To study the effects of aging processes onto the internal shear strength of multilayer geosynthetics, shear strength tests were conducted in long-term test devices, which subject the product to physical stress at high temperature in a liquid test environment until failure occurs. These test stands were originally developed for textured geomembranes. The experimental set-up of the shear test device is schematically shown in Fig. 1. By varying the test conditions, functional relationship can be established with times-to-failure, allowing extrapolation of expected service life under field conditions. For long-term shear strength testing, the GCL specimen was sandwiched between two stainless steel wedges (Fig. 2). Both wedges had a base with an area of 120  120 mm2 and a height of 48 mm, resulting in an angle of inclination of 2.5:1 or 21.81. A specimen of 230 mm length and 120 mm width was cut out of the GCL samples. The cover geotextile was carefully separated from the carrier geocomposite at both ends at a length of 50 mm. The separated geotextiles were mounted into the specimen holder as shown in Fig. 2. Firm contact of the upper and lower

Fig. 2. Specimen holder with mounted GCL specimen no. 75 removed from the test device immediately after the test end. A textured geomembrane had been fixed to the lower and a metal food grater to the upper wedges as friction surfaces.

surface of cover and carrier geotextile to the wedges was ensured by full-surface anchoring on an appropriately textured layer which was fixed on the surface of each wedge. Textured geomembranes or metal food crown graters (Fig. 3) served as a textured layer.

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stress sn on the plane of the GCL and the shear stress t in the plane of the GCL are calculated according to: sn ¼

F n F g cos a , ¼ A A

(1)

F S F g sin a , (2) ¼ A A where A is the specimen area. The following relation gives the stress on a reinforcing fiber bundle fS, t (3) fS ¼ , r

t¼

Fig. 3. Textured geomembranes or metal food graters are attached to the wedges to provide all over adhesion of the wedges to the upper and lower surfaces of the GCL specimens. The friction partners are fixed in position on the wedges by hanging them on two pins firmly screwed to the wedges. Boreholes can be seen on the upper edge.

Fig. 4. Forces and stresses on GCL specimens. a is the angle of inclination, Fn the normal force, Fg the applied force, FS the shear force and A is the sample area.

The specimen holder was housed in a controlledtemperature water bath (Tmax ¼ 80 1C). The experiments were performed at room temperature, 40, 60 and 80 1C. Tap or de-ionized water was used as a test medium. The specimens were kept under water during the whole test period. A lever mechanism was used to exert a force on the upper wedge, which is representative for loading in landfill cover applications (Fig. 1). The forces exerted on a GCL specimen are schematically shown in Fig 4. The normal

when the number of fiber bundles r per specimen area is known. In most cases, a vertical test load per base area of the wedge of 50 kPa was applied. This may be considered as a typical upper limit of loading conditions in slopes of landfill covers, although load conditions may vary considerably according to side-specific conditions. Since the base area of the wedges is 0.0144 m2, a vertical force of Fg ¼ 720 N was required. The force exerted by the weight of the thrust rod, upper wedge and mounting parts was 28 N and that by the weight of the level arm, test weight holder and displacement sensor 1 was 64 N. Since the advantage is 1:3, an additional force of (720 N28 N 64 N)/3 ¼ 209.3 N or a test weight of 21.34 kg had to be exerted. The area of the inclined plane and the specimen is about A ¼ 120  130 mm2 ¼ 0.0156 m2. This gives a compressive stress normal to the specimen plane of sn ¼ 42.9 kPa and a shear stress in the specimen plane of t ¼ 17.1 kPa (Fig. 4). The total displacement of the upper wedge and its vertical component were measured by two displacement pick-ups. Sensor 1 was mounted on top of the vertical rod to measure the vertical movement of the upper wedge. Sensor 2 detected the displacement of the upper wedge using a Bowden cable. It is possible to characterize the deformation geometry based on the two displacement records. Fig. 5 shows the two possible combinations of shear deformation and change in thickness: compression (right) and swelling (left). Some point A on the upper wedge is transferred to point B as a consequence of shear deformation and swelling or compression. The distance AB is measured by displacement sensor 2. Let D denote the positive (swelling) or negative change (compression) in specimen thickness d. The displacement s in the shear plane is then given by: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s ¼ ðABÞ2  D2 . (4) The transfer is accompanied by a vertical movement of the rod, as labeled with h and detected by sensor 1. The relationship between D, s and h is given by: s sin a  h . (5) cos a The initial length of the reinforcing fiber bundle is given by d. Prior to deformation the fibers are oriented perpendicularly to the shear plane. Due to deformation D¼

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the reinforcing fibers are stretched and tilted by the angle b. The lengths of the stretched fiber bundle are given by l. According to Eqs. (4) and (5), change of thickness and displacement in the shear plane may be calculated from the data of displacement sensors 1 and 2. The actual length of the reinforcing fiber bundles in the shear test device is given by qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l ¼ ðd þ DÞ2 þ s2 . (6) The tilt angle b of the reinforcing fiber bundles can be expressed by the cosine rule: ðABÞ2 ¼ d 2 þ l 2 þ dlcos b.

(7)

Finally, the strain e (%) can be obtained as ld . (8) d After the test end, the specimen holder was removed from the test device and a bentonite specimen was immediately taken from the dismantled GCL specimen. The water content of the bentonite was then measured according to the German standard DIN 18121 Part 1. Water adsorption (WA) was determined according to DIN 18132 (Enslin Neff) and the swell index (SI) measured based on ASTM D5890. The residual internal short-term shear strength of some GCL specimens which had passed the long-term shear strength test without failure was measured in shear box

 ¼ 100 

Fig. 5. Deformation geometry of a GCL specimen with a ¼ slope angle, b ¼ tilt angle of the reinforcing fiber bundle, d ¼ original thickness (dry thickness before being mounted), D ¼ change in thickness after deformation and l ¼ final length of fiber bundle after deformation.

tests. The test procedure observed the recommendation GDA E 3-8 of the German Geotechnical Society (Blu¨mel and Brummermann, 1994) and DIN EN ISO 12957-1. The measurements were performed at the Institute of Soil Mechanics, Foundation Engineering and Waterpower Engineering of the University of Hanover, Germany. 3. Results 3.1. Test results for commercially available products with tap water as a test medium First of all, the test results for the commercially available products GCL 2Aa, GCL 2Ab, GCL 2Ac and GCL 2Ad with tap water as a test medium are shown. It is important whether tap water (conductivity: 700 mS/cm, Na: 46 mg/l, Ca: 94 mg/l) or de-ionized water (conductivity: 1.9 mS/cm, Na: 0.032 mg/l, Ca: 0 mg/l) has been used. Tap water induced a sodium to calcium ion exchange in the bentonite layer right from the test beginning. Indeed, tap water as test liquid remained clear throughout the whole test period whereas de-ionized water rapidly became cloudy. This was due to hydrated sodium bentonite continuously pressed out and dissolved from the edges of the specimen which had been mounted on the specimen holder without lateral confinement (Fig. 2). As it will be discussed later, long-term shear behavior of GCLs in de-ionized water was completely different from the one in tap water. Table 2 shows WA and SI of bentonite specimen removed from the GCL specimens immediately after the end of the shear test. Product specification for the original sodium bentonite GCL requires WAX500% and SIX20 ml/2 g. The measured values of the tested GCLs are in agreement with typical values for calcium bentonite, which are about 200–300% WA and about 8 ml/2 g SI. In addition, the results of shear box tests on specimens which had passed the long-term shear strength test are listed. Tables 3–6 show the conditions of the long-term shear strength tests and their duration. All tests had to be terminated after long test durations either because of technical issues or for different use of the test equipment. Failure was not observed in any of the cases. Figs. 2 and 6 show, as an example, specimen no. 75 after dismantling from test device 6. This specimen, as all others, was in a

Table 2 Properties of the bentonite and results of shear box tests on GCL specimens which had passed long-term shear strength tests Sample

No. of specimen (no. of test device)

Test duration (days)

Short-term shear strength at 20 kPa normal pressure (kPa)

Bentonite water content (%)

Bentonite water adsorption (%)

Bentonite swell index (ml/2 g)

GCL GCL GCL GCL GCL GCL

11 (10u) 19 (8u) 18 (5u) 61 9 39

713 713 1125 929 927 927

62 70 60 66 26 31

120 119 107 109 104 110

275 350 315 300 230 340

10 10 10 10 8 10

2Aa 2Aa 2Aa 2Ab 2Ad 2Ad

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Table 3 Time elapsed and test conditions for long-term shear strength tests on GCL 2Aa No. of specimen (no. of test device)

Test temperature (1C)

Test load (kPa)

Angle of test wedge (1)

Friction partner lower wedge

Friction partner upper wedge

Test duration (days)

Comments

18 33 43 51 75 25 53 57 11 19

80 80 80 80 80 80 80 80 60 60

46 50 50 50 50 50 50 50 50 50

21.8 21.8 21.8 21.8 21.8 21.8 21.8 21.8 21.8 21.8

Textured GM Textured GM Textured GM Textured GM Textured GM Friction plate Textured GM Textured GM Textured GM Textured GM

Friction Friction Friction Friction Friction Friction Friction Friction Friction Friction

1125 323 1351 637 591 553 415 415 713 713

Terminated Terminated Terminated Terminated Terminated Terminated Terminated Terminated Terminated Terminated

(5u) (16) (2u) (9) (6) (11+21)

(10u) (8u)

plate plate plate plate plate plate plate plate plate plate

no no no no no no no no no no

failure failure failure failure failure failure failure failure failure failure

GM: geomembrane.

Table 4 Time elapsed and test conditions for long-term shear strength tests on GCL 2Ab No. of specimen (no. of test device)

Test temperature (1C)

Test load (kPa)

Angle of test wedge (1)

Friction partner lower wedge

Friction partner upper wedge

Test duration (days)

Comments

61 (1u)

80 80

50 50

21.8 21.8

Textured GM Friction plate

Friction plate Friction plate

540 276

(2u)

80

50

21.8

Friction plate

Friction plate

276

(3u)

80

50

21.8

Friction plate

Friction plate

276

68 (17+27)

80

50

21.8

Friction plate

Friction plate

170

74 (20+30)

80

50

21.8

Friction plate

Friction plate

170

Terminated no failure Specimen pre-aged by 202 days aging at 80 1C, progressinga Specimen pre-aged by 202 days aging at 80 1C, progressinga Specimen pre-aged by 202 days aging at 80 1C, progressinga Specimen pre-aged by 505 days aging at 80 1C, progressinga Specimen pre-aged by 505 days aging at 80 1C, progressinga

a

oven oven oven oven oven

Test duration time up to April 19, 2007.

Table 5 Time elapsed and test conditions for long-term shear strength tests on GCL 2Ac No. of specimen (no. of test device)

Test temperature (1C)

Test load (kPa)

Angle of test wedge (1)

Friction partner lower wedge

Friction partner upper wedge

Test duration (days)

Comments

50 (12+22) 48 (13+23)

80 80

50 50

21.8 21.8

Friction plate Friction plate

Friction plate Friction plate

482 482

Progressinga Progressinga

a

Test duration time up to April 19, 2007.

perfect condition by all appearances. Specimens numbered 9, 39, 53, 57 and 61 were placed into shear test devices, which were installed in the CGL manufacturer laboratory. Also, these test specimens were found to be still intact when removed. Short-term shear strength measured by shear box tests was still comparable to the initial strength of the samples GCL 2Aa (Table 2). A considerable reduction in shear strength compared to typical initial values has occurred for samples GCL 2Ad, where a woven geotextile alone acts as carrier geotextile.

Since no failure has occurred, the average test duration can only serve as a lower limit for possible time-to-failure, which in actual fact would have to be considerably longer. The longest test on GCL 2Aa had been running for more than 312 years and the geometrical mean value from five tests with the longest running time is decidedly more than 2 years: 868 days or 20,800 h. Fig. 7 shows the displacement AB (approximately equal to the shear displacement since D5s) and the change in thickness D of GCL 2Aa specimen no. 25 (device 11+21),

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Table 6 Time elapsed and test conditions for long-term shear strength tests on GCL 2Ad No. of specimen (no. of test device)

Test temperature (1C)

Test load (kPa)

Angle of test wedge (1)

Friction partner lower wedge

Friction partner upper wedge

Test duration (days)

Comments

9 39

80 80

50 50

21.8 21.8

Textured GM Textured GM

Friction plate Friction plate

540 540

Terminated no failure Terminated no failure

GM: geomembranes.

Fig. 6. View on GCL specimen no. 75, which was dismantled after 591 days of testing time. The imprints are due to the textured geomembrane.

calculated by Eq. (5). The increase in thickness by swelling under the test load is about 1.5 mm. Since the initial thickness of the dry specimen was about 7 mm, one obtains 10 mm for the length l of the elongated fiber bundle from Eq. (6). The test results demonstrate that there has been no detrimental effect on the friction parameters between the textured geomembrane and the GCLs due to aging effects under the current test conditions. In addition to shear strength testing, the oxidative resistance of the geosynthetic components was investigated by oven aging measurements. The test results and details of the experiments on the nonwoven geotextiles are described elsewhere (Mueller et al., 2003). The nonwoven geotextiles made of PP-resins PP1 and PP2 showed high resistance against oxidative degradation (Thomas, 2002). The woven geotextile used in the carrier geocomposite was sensitive to some extent to thermal oxidative degradation. Fig. 8 shows the reduction in relative tensile strength (actual tensile strength related to initial tensile strength) during oven aging at 80 1C. For this, the slit film specimens taken from GCL samples were hung up on a grid in a thermostatically regulated Heraeus UT 6200 oven. The aging temperature was achieved and maintained by controlled heating of the working chamber walls. The air exchange took place via two open air channels (3 cm diameter) on top and bottom of the back wall. An air ventilation fan in the ceiling of the working chamber forced internal circulation. Tensile

Fig. 7. Change in thickness and shear elongation (sEAB) of GCL specimens, calculated by Eq. (5).

strength of slit film specimens taken from the GCL sample was considerably lower than that of specimens taken from the original fabric, probably due to damage caused by the needle-punching process. A further reduction in tensile strength occurred during air oven aging at 80 1C. There was no indication of any reduction in tensile strength of the slit films of the woven geotextile made of PP3 from specimens GCL 2Aa during the long-term shear strength testing in hot water. To study the effects of oxidative degradation of the woven geotextile on GCL shear strength, samples of GCL 2Ab (resin PP4) were pre-aged in a thermostatically regulated Heraeus T 6760 oven at 80 1C (without internal air circulation). After 505 days of aging time the tensile strength of the slit film specimens extracted from the GCL was only about one-fourth (473 N) of the initial value (1474 N). These pre-aged GCL specimens taken from the oven after 202 and 505 days were installed into the shear strength test device. Table 4 shows the test conditions and test duration for these specimens. 3.2. Test results with de-ionized water as test medium and failure modes Samples GCL 1A, GCL 1B, GCL 2B and GCL 2Aa, i.e. samples made of PE and PP, each thermal fixed and not

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Fig. 8. Change of tensile strength relative to the initial strength of slit film woven specimens used in the carrier geocomposite during air oven aging with internal air circulation at 80 1C. See Table 1 for specimen designation.

Fig. 9. Length of fiber bundles l at failure as a function of the associated time-to-failure and temperature.

thermal fixed, were tested using de-ionized water as a test medium. Under these test conditions, ion exchange was limited and the sodium bentonite exhibited a high water content and plasticity throughout the test duration. As a result sodium bentonite was slowly squeezed out of the mounted specimen accompanied with a large displacement in the shear plane. Thereby, the reinforcing fiber bundles were considerably strained with small tilt angles (Fig. 9). Failure, i.e. sliding of the upper wedge, occurred regularly under this condition. Figs. 10 and 11 show the cumulative frequency distribution of the lognormal distributed timesto-failure measured for specimens GCL 1A and GCL 2Aa at 80 1C. Different failure modes were observed for the two manufacturing methods A and B. Fig. 12 shows a specimen holder with specimen GCL 2Aa, i.e. a specimen with high

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Fig. 10. Cumulative frequency distribution of times-to-failure measured for GCL 1A at 80 1C in de-ionized water.

Fig. 11. Cumulative frequency distribution of times-to-failure measured for GCL 2Aa at 80 1C in de-ionized water.

anchor strength by thermal treatment. The wedges were dismantled from the test device immediately after failure had occurred. Bentonite was squeezed out at the specimen edges. In the central bentonite filled specimen area, the reinforcing fiber bundles were ruptured near their anchoring base. Only the fiber bundles on the left and right specimen margin were elongated and partially pulled out of the cover nonwoven geotextile. This type of failure mode was typical for all thermally treated GCLs with high anchoring strength made of PE as well as of PP. In the case of thermal fixed GCLs the ends of needlepunched fiber bundles standing out of the carrier geocomposite are partially molten to lumps and fixed on the outer side of the geocomposite due to the thermal treatment. A reinforcing fiber bundle extends back from such a lump through the bentonite layer into the fiber tangle of the cover nonwoven geotextile. Fig. 13 gives a closer look on the ruptured fiber bundle fragments which are jutting out

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Fig. 14. Fiber bundle fragment from the carrier geocomposite of a failed, thermal fixed specimen which corresponds to the fiber bundle fragments of Fig. 13. Fragments of fibers extend from the lump of molten fibers which was formed during thermal treatment, see Fig. 20.

Fig. 12. Specimen holder (specimen GCL 2Aa, thermal fixed, de-ionized water) dismantled from the long-term shear test device after failure due to sliding of the upper wedge.

Fig. 13. Inner side of the cover nonwoven geotextile from a failed, thermal fixed specimen after removal of bentonite and cleaning. The network of fragments of fractured needle-punched fiber bundles is seen which are jutting out of the cover nonwoven geotextile (crew cut), see Fig. 20.

from the cover nonwoven geotextile. Fig. 14 shows a corresponding fragment which was dissected from the carrier geocomposite. Most of the fibers ruptured in the region between the bentonite layer and the carrier nonwoven geotextile. However, scanning electron micro-

scopy (SEM) reveals that rupture of the fibers seems to have occurred at the lump base, too (Fig. 15a and b). Not thermal fixed specimens showed a completely different failure mode. To be not thermal fixed means that the fiber bundles from the cover nonwoven geotextile are simply needle-punched through the bentonite layer and through the carrier geocomposite (nonwoven and woven geotextile). The bundle end is loosely entangled in the thin nonwoven geotextile, which covers on the outer side of the woven geotextile. Fig. 16 shows a specimen holder with the not thermal fixed specimen GCL 2B, which was dismantled from the test device after failure. Again, bentonite had been squeezed out at the specimen edges. During shear strength testing the fiber bundles had been slowly pulled out of the carrier geocomposite. Finally, failure occurred when the pull-out resistance of the imbedded fiber bundles fell below the pull-out force. The carrier geocomposites of failed specimens were completely intact without any residues from needle-punched fiber bundles. Fig. 17 shows the cover geotextile nonwoven with the stretch-out, intact fiber bundles from the needle-punching process. The times-to-failure, which were observed in tests with de-ionized water as a test medium, strongly depended on resin and test temperatures. Figs. 18 and 19 show Arrhenius plots of times-to-failure versus reciprocal absolute temperature. Each data point is the geometric average of various single measurements (Figs. 10 and 11). Thermal fixed GCL 1A and GCL 2Aa specimen failed much faster than not thermal fixed GCL 1B and GCL 2B. For the thermal fixed specimens, which showed the failure mode of rupturing fiber bundles, a linear relation between reciprocal absolute temperature and the logarithm of times-to-failure was found. The activation energy is 5677 kJ/mol for GCL 1A, and 80717 kJ/mol for GCL 2Aa. The times-to-failure at the selected test conditions are so short that embrittlement due to oxidative degradation of the fibers is not relevant. The fiber breaking mechanism

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Fig. 15. (a) SEM picture of a lump from a failed specimen GCL 1A. (b) SEM picture of a ruptured fiber from a failed specimen GCL 2Aa. Rupture had occurred at the lump base of the fiber.

Fig. 17. Cover GTnw from a failed, not thermal fixed specimen (GCL 2B) still filled with bentonite. The fiber bundles from needle-punching had been completely pulled out of the carrier geocomposite.

Fig. 16. Specimen holder with specimen GCL 2B (not thermal fixed, deionized water) dismantled from the long-term shear test device after failure due to sliding of the upper wedge.

seems to be accelerated by mixing surfactants into the de-ionized water bath. Table 8 gives a comparison of shear strength evaluated by long-term tests with the results of short-term peel tests based on DIN EN ISO 10319 (von Maubeuge and Lucas, 2002). The short-term shear strength of GCLs is expressed as the maximum shear force (shear strength) or maximum peeling force (peel strength) that can be observed in shortterm shear or peel tests. Their long-term shear strength can

be characterized by the shear stress withstood by 95% of the samples tested over a given test duration at a specified temperature. It can also be characterized by the geometric mean of the time-to-failure in long-term shear testing at specified shear stress and temperature. The short-term peel test showed that GCLs manufactured with method A (thermal fixed) have higher peel strength compared to method B (not thermal fixed). Obviously a higher force is necessary to break the fixed fiber bundles than to disentangle them at high test speed in a short-term peel test. Contrary to this, the times-to-failure of GCL 1B clearly exceeded those of GCL 1A at all temperatures in the long-term test. The same result holds for GCL 2Aa compared to GCL 2B. At a constant shear force level long-term brittle fracture of the fixed fiber bundles develops faster than their disentanglement without fixing. The fiber bundle length l at failure was calculated by Eq. (6) (with the approximation ABEs) and plotted as a

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4. Discussion 4.1. Long-term shear behavior

Fig. 18. Arrhenius diagram of shear test results for GCL 1A and 1B. Arrows indicate mean elapsed times which include times from tests terminated without failure.

Fig. 19. Arrhenius diagram of shear test results for GCL 2Aa and 2B. Arrows indicate mean elapsed times which includes times of tests terminated without failure. The line fits to the data of GCL 2Aa, deionized water.

function of times-to-failure (Fig. 9). The figure shows that the fiber bundle length decreases with times-to-failure or increases with testing temperature. For thermal fixed specimens the fiber length at failure seems to be much lower than for not thermal fixed specimens. The calculation of the fiber bundle length according to Eq. (6) includes the stretching and tilting of the reinforcing fiber bundles but also the stretching of the cover nonwoven geotextile itself. Long-term shear tests without water were performed at room temperature. In this case, the displacement was very slow: no movement was observed for GCL 1B after 113 days and only 1 mm displacement of the upper wedges for GCL 1A after 63 days. The internal friction angle of the dry sodium bentonite is obviously high enough to prevent shear movement of the specimens.

Results of short-term tests (e.g. peel tests or shear box tests) are currently used for designing against internal shear failure of GCLs. However, the long-term slope stability depends on the ability of the reinforcement of connecting fiber bundles to sustain stress over a very long time period. Therefore, the loading conditions of the reinforcing fiber bundles and their aging behavior under such conditions in the long run have to be taken into account. When long-term shear strength of GCLs is tested in deionized water, ion exchange in the sodium bentonite layer is prevented. The hydrated sodium bentonite layer of the GCL has essentially no intrinsic shear strength and the shear stress is fully carried by the highly strained fiber bundles. Under this condition brittle failure of the thermally treated GCLs and ductile failure due to slow disentanglement of the fibers of the not thermal fixed GCLs was observed. Brittle failure is characterized by fracture of the fibers especially at their anchor points in the carrier geocomposite, ductile failure by the complete disentanglement of the fiber bundle from the carrier geocomposite. Thermal fixing has the advantage to increase the internal shear strength of the GCL. Yet, it has to be shown by long-term shear strength testing, that there are no detrimental effects of the thermal treatment, which can make the material susceptible to brittle failure. With time, both failure mechanisms result in a total loss of internal shear strength. Fig. 20 shows a schematic illustration of the two failure modes. An Arrhenius-like temperature dependence was proven for the brittle failure mode and activation energies could be attributed to this failure mechanism. Times-to-failure of GCLs with PE geotextiles are shorter than with PP geotextiles for both failure modes. In the case of the failure mode of disentanglement of fiber bundles, this can be due to the lower melting temperature of PE. PE fibers are softer at 80 1C than the PP fibers and this could make the disentanglement easier. The higher stress crack sensitivity of the PE fiber resin compared with the PP fiber resin should give shorter times-to-failure for the mode of brittle rupture of the thermal fixed fiber bundles. A totally different long-term shear behavior of the GCL was observed when tap water was used instead of deionized water as a test medium. Under this test condition the ion-exchanged calcium bentonite reinforced by thermal fixed fiber bundles formed a compact layer which had a shear resistance above the applied test shear stress (Fig. 2). The fiber bundles seemed to be only elongated and stressed below their failure limit. Even after more than 3 years of test duration the dismantled specimens were completely intact by appearance. In shear box tests, specimens GCL 2Aa showed short-term internal shear strengths comparable to the initial values. A reduction in internal shear

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Fig. 20. Schematic illustration of failure modes of GCL specimens in long-term shear strength testing with de-ionized water as a test medium.

strength occurred only for specimen GCL 2Ad, where a woven geotextile acts alone as carrier. The effects of sodium to calcium ion exchange on the shear properties of the sodium bentonite, which was used for the GCL samples was studied by Weber (2006). Layers of sodium bentonite encased in geotextile bags were loaded with 30 kPa and immersed for 7 days at 80 1C in de-ionized and tap water and in a 3.33 mol/l CaCl2 solution. Table 7 shows the results. Ion exchange was small after 7 days of immersion in tap water, but developed to a large extent in the CaCl2 solution. The angles of internal friction of the ion-exchanged bentonite specimens immersed in CaCl2 solution were indeed considerably higher than the friction angle of the specimen in de-ionized water. 4.2. Estimates of service life Results of long-term shear strength tests may be used to estimate service life by taking into account that the time-tofailure strongly depends on temperature. Fracture initiation may first progress unseen and shear strength failure occur only after a certain accumulation of damage. A change in property linked to such processes may often be approximated by an exponential time dependence whereby an Arrhenius relation between the velocity constant k of the exponential decay and the temperature is observed: k ¼ k0 eE=RT .

(9)

E is the apparent activation energy of the process to be assessed and R ¼ 8.314 J/(mol K) is the general gas constant. Whether such a phenomenological description can be applied or not can be seen with the help of the

Table 7 Bentonite properties due to sodium to calcium ion exchange after 7 days of immersion at 80 1C Bentonite properties

De-ionized water

Tap water

3.33 mol/l CaCl2

Water content (%) Water adsorption (%) Liquid limit (%) Cation exchange capacity (mmol(eq)/100 g) Angle of internal friction (1)

196 686 566 77

187 700 530 71

43 213 73 50

a

13

18; 7a

25; 33a

Two specimens.

Arrhenius diagram (Fig. 19). It requires that the logarithm of times-to-failure measured at different temperatures varies linearly when plotted over the reciprocal (absolute) test temperature. The slope of the line determines the activation energy. Extrapolation of a service life tF at a field temperature TF from time-to-failure tS at a test temperature TS may be calculated as: tF ¼ tS eðE=RÞ½1=T F 1=T S  .

(10)

In agreement with these considerations, 80717 kJ/mol or 960072000 K was estimated for the activation energy of the brittle failure mode found for needle-bonded and thermally treated GCL 2Aa when tested in de-ionized water. Assuming that the underlying failure mechanism may become relevant in the long run for ion-exchanged GCL as well, their service life may be estimated for commercially available products from the test duration times in tap water. The mean value of the five longest test

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duration times for GCL 2Aa in the long-term shear strength test (80 1C and 50 kPa) employing tap water as a test medium is more than 2 years (Table 3). With an apparent activation energy of 7600 K (mean value minus standard deviation), which is associated with the fiber breaking failure mode, a lower limit of functional durability of at least 250 years at TF ¼ 15 1C can be derived from Eq. (10). Various results from excavation, field and laboratory experiments have shown, that a total exchange of sodium by calcium ions will take place in a sodium bentonite GCL within a few years, when hydrated under field conditions (Egloffstein, 2001; Blu¨mel et al., 2002a, b; Bouazza et al., 2007). This period of time is short compared with an expected functional durability of ion-exchanged GCLs of far more than 100 years, as discussed above, and short compared with a hypothetical Arrhenius-extrapolated lifetime without ion exchange of about 30 years. Therefore, it is justifiable only to consider the results of tests carried out using tap water in the assessment of field performance. Estimations based on exponential relations are naturally subject to large uncertainties. Even when taking into account these uncertainties, according to the data at hand, static functional durability of GCL 2Aa under typical geotechnical application temperatures and loading conditions should be considerably more than 100 years, a lower service life limit, which is set by German regulations. The issue of how various on-site conditions in comparison to laboratory conditions are to be evaluated still remains to be discussed. Previous investigations into longterm shear strength of geosynthetics repeatedly showed that the times-to-failure sensitively depend on the selected polymer materials, production methods, properties of the product and the test conditions (Mu¨ller et al., 2004). The test results given here may therefore not be transferred to other products or application conditions which differ from the one described here. In the long-term shear strength test, the specimen is stressed between two completely even friction surfaces with a test load of 50 kPa. This corresponds at a unit weight of soil material of 20 kN/m3 to a quite large reclamation layer thickness of 2.5 m. With inclination of the sloping test plane of 2.5:1, shear stress is 17.1 kPa and compressive stress 42.9 kPa. The internal shear strength of bentonite decreases with increasing water content. As the test was carried out under water, the bentonite had a water content of distinctly more than 100% (Table 2) which is higher than the range of water content expected under field conditions. Clearly, the layer of calcium bentonite and thermal fixed fiber bundles, from which the internal shear resistance results, is stressed to a maximum in the tests. Compared to these laboratory conditions, there are still safety reserves under site conditions where the subgrade layer beneath and the covering layer above the GCL will not be even, shear stress is lower and bentonite scarcely has such a high water content. However, it was not possible to quantify these reserves.

4.3. Effects of oxidation The most significant aging mechanism of polyolefin products is embrittlement brought about by oxidative degradation. A detailed long-term investigation on PP and PE nonwoven geotextile showed that the conditions for oxidation have a very great impact on long-term behavior (Thomas, 2002; Mueller et al., 2003). On one hand, the potential failure mechanisms in the relatively complex structure of multilayer geosynthetics under permanent and combined impact of compression and shear could only be shown in long-term shear strength tests. On the other hand, it was shown in Mueller et al. (2003) that immersion in an 80 1C water bath, as carried out in these long-term tests, realizes only a weak oxidative stress for nonwoven geotextiles compared, for example, with testing in oven aging with forced air circulation. This applies especially for hindered amine stabilized materials. Oxidative stress onsite will lie somewhere between these two experimental conditions, namely when installed constantly under water or constantly surrounded by fresh air, whereby being constantly surrounded by air would be quite certainly a very unusual and extreme case and the application condition will probably tend to be similar to immersion in water. Nonetheless, one should check oxidation stability under critical oxidation conditions. Investigations had been made on the nonwoven geotextile PP1 and PP2 regarding thermal oxidative degradation in oven aging (Thomas, 2002; Mueller et al., 2003). Such investigations are part of the tests within BAM certifications for geotextile components for protective layers of geomembranes and BAM approval for geosynthetic drains, where these nonwoven geotextiles are used as filter and protection layers (Mu¨ller, 2006, 2007). The nonwoven geotextiles PP1 and PP2 were approved by BAM as the material showed high oxidation stability even under critical conditions. Complementary investigations were performed for the woven geotextile used in the carrier geocomposite (Fig. 8). The woven geotextiles PP4 and PP5 are obviously poorly stabilized and a rapid reduction in tensile strength is observed. Such products should not be used in GCLs for long-term application. Reduction in mechanical strength of PP3 proceeds more slowly. The time dependence of the degradation of mechanical strength is similar to those observed for hindered amine stabilized PP materials (Gijsman, 1994; Zweifel, 2001). However, no information about the stabilizer package is available. In GCL 2Aa, GCL 2Ab and GCL 2Ac the carrier geocomposite is a composite formed by connecting the woven and nonwoven geotextiles by needle-punching. The results of the longterm shear strength test on such GCLs, which were preaged by air oven aging, showed that the reduction in tensile strength of the woven geotextile made of PP3 does not induce shear failure in the test device (Table 4). The tensile strength of the woven geotextile, which is part of the carrier geocomposite, seems to be of minor importance for

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internal long-term shear strength of the GCL. However, the long-term performance of sample GCL 2Ad, where a woven geotextile acts alone as carrier geotextile, seems questionable with respect to critical oxidation conditions in view of these results and the reduction in initial tensile strength, which was observed for long-term tested specimens GCL 2Ad. It seems to be plausible, that the reduction in initial tensile strength, which was observed for long-term tested specimens GCL 2Ad, has to be attributed to oxidative degradation of mechanical strength of the carrier woven geotextile made of PP5. However, no data of shear strength tests on pre-aged GCL 2Ad are available so far. 4.4. Short-term versus long-term shear strength Short-term tests involve simple experimental procedures and are time-saving. Therefore, they are frequently used as dimensioning data—modified by safety factors as appropriate—in geotechnical structural design. For all building materials, short-term (i.e. high-speed) strength can differ significantly from the actual long-term strength observed under field conditions. In addition, all building materials are subject over time to aging processes (corrosion, stress cracking, morphological changes, oxidation, etc.), which affect physical strength. The resultant strength reduction varies greatly over time and cannot be reliably characterized by safety factors (Mu¨ller, 2006, 2007). Aside from shoddy workmanship, insufficient differentiation between short-term and long-term properties of construction materials is the single most common cause of structural failure with often dramatic consequences. Shear strength of geosynthetic products is normally determined in shear box tests (von Maubeuge and Lucas, 2002). In this procedure, the upper and lower surfaces of the tested product are fixed in place on the respective halves of the shear box apparatus, which are compressed by a defined normal load and moved apart at constant speed. The resultant shear force is recorded as a function of displacement. This allows a determination of the peak shear force. A similar data plot can be obtained for peel strength, which is measured by peeling the samples in a wide width tensile test. Peel strength is closely related to shear strength, but much easier to measure. For this reason, products are optimized with respect to short-term peel strength (von Maubeuge and Lucas, 2002). Short-term behavior of GCLs as reflected by peel strength is in clear contrast to the long-term behavior. Table 8 compares mean peel strengths from peel tests and mean relative times-to-failure from long-term shear tests of different GCLs specimen. To facilitate comparison, all data are expressed as relative values, i.e. multiples of the smallest value in each data group. The data clearly show that high short-term strength is unrelated to high times-to-failure under long-term shear stress. A design of cover system with GCLs on steep slopes is therefore questionable when solely based on the analysis of the shortterm behavior, i.e. either on peel strength or short-term

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Table 8 Relative peel strength from short-term peel test and relative times-tofailure from long-term shear test of different GCL products Product number

Relative peel strength

Relative time to failure

GCL GCL GCL GCL

3.8 3.6 2 1

1 4.4 24 4270

1A 1B 2A 2B

shear strength. Long-term shear testing is an essential prerequisite in the assessment of shear behavior of multilayer geosynthetics consisting of several structural or material components. 5. Conclusion A new test method, originally developed for textured geomembranes, was applied to measure long-term internal shear strength of various types of needle-punched GCLs. The test conditions were highly demanding—high temperatures, compressive stress of 43 kPa, shear stress of 17 kPa and continuous submersion in water—and selected with the specific purpose of accelerating potential failure mechanisms. The following findings have been gained from these experiments. The long-term shear behavior of the GCLs strongly depended on the test medium. Tap water is essential to ensure sodium to calcium ion exchange and thus avoiding the large straining and failure of the reinforcing fiber bundles. The layer of ion-exchanged bentonite and reinforcing fiber bundles has an intrinsic long-term shear resistance which is well above the applied shear force of 17 kPa at 45 kPa normal pressure. No failure was observed even after 3 years of testing at 80 1C when tap water was used. Since under normal field conditions ion exchange will occur within a few years, shear strength of the tested GCLs will be high enough for typical landfill cover conditions in the long run. With de-ionized water the sodium bentonite layer showed high water content and plasticity and a large elongation of the connecting fiber bundles was induced. Under this condition failure was imminent and the timesto-failure strongly depended on fiber resin, product design and temperature. Two failure modes could be identified depending on the manufacturing process. GCLs with thermal fixed fiber bundles due to thermal treatment showed brittle failure by rupture of the fiber bundles near the anchoring points in the carrier geocomposite. GCLs, which were not thermal fixed, showed a complete disentanglement of the fiber bundle form the carrier geocomposite. The short-term bonding strength, which was measured in peel tests, is unrelated to the times-to-failure from the longterm shear test. Surprisingly, the GCL types with higher bonding strength had shorter mean times-to-failure in the

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long-term shear test. Therefore, a high bonding strength in peel test is not a reliable indicator of high shear strength in the long run. In (Mueller et al., 2003) it was shown that for nonwoven PP and PE geotextiles air oven aging is more critical than immersion in water. This is, because PP geotextiles are often stabilized by hindered amines, which deplete more rapidly in the oven than in the water bath. In addition, the structural stabilization of the highly oriented fibers seems to be more effective under the oxidation condition in water than in air. To assess the long-term performance of the GCL, one should test the oxidation stability of the geotextiles under air oven conditions in addition to longterm shear strength tests. The reduction in internal strength of specimen GCL 2Ad after long-term shear strength testing in tap water is probably due to the pure oxidation stability of the woven carrier geotextile, which was revealed in the air oven aging. Acknowledgments The presented results originated from a BAM research project, which was supported with a (moderate) funding and technical assistance by NAUE GmbH & Co. KG. The authors would like to thank Professor W. Blu¨mel and M. Heinemann, Institute of Soil Mechanics, Foundation Engineering and Waterpower Engineering of the University of Hanover, Germany, who performed the shear box tests. References Blu¨mel, W., Brummermann, K., 1994. Reibung zwischen Geokunststoffen und Erdstoffen in Deponiedichtungen. Mu¨ll und Abfall 26 (5), 242–259. Blu¨mel, W., Mu¨ller-Kirchenbauer, H., Ehrenberg, H., von Maubeuge, K, 2002a. Langzeituntersuchungen zur Wasserdurchla¨ssigkeit von Bentonitmatten in Lysimetern. In: Egloffstein, T.A., Burkhardt, G., Czurda, K. (Eds.), Oberfla¨chenabdichtungen von Deponien und Altlasten 2002. Erich Schmidt Verlag, Berlin, pp. 183–198. Blu¨mel, W., Mu¨ller-Kirchenbauer, H., Markwardt, N., 2002b. Lysimeteruntersuchungen zu Wasserdurchla¨ssigkeit und -haushalt an Deponieabdichtungssystemen mit Bentonitmatten. Geotechnik 25 (4), 261–270. Bonaparte, R., Daniel, D.E., Koerner, R.M., 2002. Assessment and recommendations for improving the performance of waste containment systems. US EPA, Cincinnati, OH. Bouazza, A. (Ed.), 2000. Special issue on geosynthetic clay liners. Geotextiles and Geomembranes 18 (2–4). Elsevier Science Ltd. Bouazza, A., Jefferis, S., Vangpaisal, T., 2007. Investigation of the effects and degree of calcium exchange on the Atterberg limits and swelling of geosynthetic clay liners when subjected to wet-dry cycles. Geotextiles and Geomembranes 25 (3), 170–185. Dickinson, S., Brachmann, R.W.I., 2006. Deformation of a geosynthetic clay liner beneath a geomembrane wrinkle and coarse gravel. Geotextiles and Geomembranes 24 (5), 285–298. Egloffstein, T.A., 2001. Natural bentonites—influence of the ion exchange and partial desiccation on permeability and self-healing capacity of bentonites used in GCLs. Geotextiles and Geomembranes 19, 427–444. Gijsman, P., 1994. The long-term stability of polyolefins. Thesis. Technical University Eindhoven, Eindhoven, The Netherlands.

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