Influence of matric suction on geotextile reinforcement-marginal soil interface strength

Geotextiles and Geomembranes 42 (2014) 139e153

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Influence of matric suction on geotextile reinforcement-marginal soil interface strength Danial Esmaili 1, Kianoosh Hatami*, Gerald A. Miller 2 School of Civil Engineering and Environmental Science, University of Oklahoma, 202 W. Boyd Street, Room 334, Norman, OK 73019, USA

a r t i c l e i n f o

a b s t r a c t

Article history: Received 8 June 2013 Received in revised form 11 January 2014 Accepted 27 January 2014 Available online 28 February 2014

This paper presents descriptions and results of multi-scale pullout and interface shear tests on a woven polypropylene (PP) geotextile reinforcement material in a marginal quality soil. A main objective of these tests was to develop a moisture reduction factor (MRF) for the pullout resistance equation in the currently available design guidelines. The tests were carried out at different overburden pressure and gravimetric water content (GWC) values. The differences in the soil-geotextile interface strength among the cases with different GWC values were used to determine the corresponding MRF values. Results of the study indicate that the reinforcement interface strength and pullout resistance could decrease significantly as a result of the loss in the matric suction (e.g. by 42% between the cases of 2% dry and 2% wet of the soil optimum moisture content). It is concluded that wetting of the soil-geotextile interface during construction or service life of a reinforced soil structure can measurably reduce the interface strength and pullout resistance of the geotextile reinforcement which needs to be accounted for in design. Results of the study will be also useful to estimate the difference in the pullout capacity of geotextile reinforcement in a marginal soil when placed at different GWC values during construction. The methodology described in the paper could be used to expand the database of MRF results to include a wider range of soil types and geotextile reinforcement for practical applications. Ó 2014 Elsevier Ltd. All rights reserved.

Keywords: Geotextiles Moisture reduction factor Marginal soils Soil matric suction Interface strength Pullout resistance

1. Introduction Transportation agencies worldwide are faced with the persistent problem of landslides and slope failures along highways, railroads and other transportation-related infrastructure. Repairs and maintenance work due to these failures are extremely costly. In Oklahoma, many of these failures occur in the eastern and central parts of the state due to higher topography and poor soil type (Hatami et al., 2010a,b, 2011). A recent example of these failures is a landslide on the US Route 62 in Chickasha, Oklahoma (Fig. 1). An ideal solution for the construction or repair of highway slopes and embankments is to use coarse-grained, free-draining soils to stabilize these structures as recommended by design guidelines and specifications for Reinforced Soil Slopes (RSS) and Mechanically Stabilized Earth (MSE) structures in North America (e.g. Elias et al., 2001; Berg et al., 2009). However, coarse-grained soils are not commonly available in many parts of the world.

* Corresponding author. Tel.: þ1 (405) 325 3674; fax: þ1 (405) 325 4217. E-mail addresses: [email protected] (D. Esmaili), [email protected] (K. Hatami), [email protected] (G.A. Miller). 1 Tel.: þ1 (405) 325 5911; fax: þ1 (405) 325 4217. 2 Tel.: þ1 (405) 325 4253; fax: þ1 (405) 325 4217. 0266-1144/$ e see front matter Ó 2014 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.geotexmem.2014.01.005

Consequently, the fill material and transportation costs can be prohibitive depending on the location of the borrow source for the high-quality soil. One possible solution in such cases is to use locally available soils as construction materials because they would require significantly less material transportation, fuel consumption and generated pollution as compared to using high-quality offsite soils. It has been estimated that fuel costs could constitute as much as 20% of the total transportation costs of high-quality soils (Ou et al., 1982). However, locally available soils for the construction of reinforced slopes in many parts of the world are of marginal quality (e.g. soils with more than 15% fines). Geosynthetic reinforcement is a wellestablished and cost effective technology for the construction and repair of slopes and embankments (e.g. Berg et al., 2009). For instance, it has been reported that reinforcing marginal soils could help reduce the cost of fill material by as much as 60% (Keller, 1995). However, proper drainage and adequate soil-reinforcement interface strength are essential elements for reinforced soil structures built with marginal soils in order to provide safe and satisfactory performance during their service life. Mechanical response of marginal soils and that of their interface with geosynthetic reinforcement are complex and may include strain softening, excessive deformation and loss of strength as a result of wetting (Zheng et al., 2013). Loss of

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Fig. 1. Failed slope of a highway embankment in Chickasha, OK. Note: the height and inclination angle of the slope are z12 m and 17, respectively.

strength due to increase in the GWC can be especially crucial at the soil-reinforcement interface because depending on the type of geosynthetic employed (e.g. geotextiles) this interface can act as a potential slip plane. This can be understood by noting the fact that interface friction coefficient values are typically less than unity for many types of geotextiles (e.g. Koerner, 2005). Moreover, failure of reinforced earthen structures that are built with marginal soils and lack a proper drainage system may simply occur in the form of significant wetting-induced deformations as opposed to complete collapse. As a result, there have been documented cases of serviceability problems and failures of these structures related to the use of marginal soils without adequate care in their design and/or construction (e.g. Zornberg and Mitchell, 1994; Mitchell and Zornberg, 1995; Christopher et al., 1998; Koerner et al., 2005; Sandri, 2005). Hamid and Miller (2009) studied the shearing behavior of an unsaturated low-plasticity fine-grained soil using a modified direct shear test apparatus in which the matric suction of the soil specimen was controlled. Their results showed that the matric suction contributed to the peak shear strength of unsaturated interfaces but did not significantly influence their post-peak shear strength. However, variations of the net normal stress affected both the peak and post-peak shear strength values. Liu et al. (2009) carried out a series of direct shear tests to study the interface shear strength of geogrids and geotextile embedded in sand and gravel. The test results showed that the shear strength of the soil-geotextile interface was 0.7 and 0.85 of the soil shear strength for Ottawa sand and gravel, respectively. Results also indicated that the shear strength of soil-geogrid interfaces was generally higher than that of soil-geotextile interfaces. The soilgeogrid interface shear strength was found to vary between 0.89 and 1.01 of the soil shear strength for the types of geogrids tested. Anubhav (2010) conducted a series of direct shear tests to examine the shear stress-displacement behavior of sand-geotextile interfaces. The results indicated that the peak shear strength of the interface between the sand and a coarse-textured geotextile was significantly higher (i.e. up to 35%) than that for the interface between the sand and a fine-textured geotextile. The results also showed that the shear displacement at peak shear strength increased with overburden pressure. In this study, it is postulated that adequate internal drainage capacity exists in the reinforced soil structure to prevent the development of positive pore water pressure in the soil. However, seasonal variations of the moisture content due to precipitation or subsurface water infiltration could still result in significant changes in matric suction during the service life of the structure in the absence of a proper global drainage system in the structure, or if for instance, the existing drainage system is compromised as a result of excessive clogging. A conservative design approach for slopes and

embankments is to assume that the embankment soil is fully saturated. However, this is not an ideal design approach for the following reasons: the soil properties from tests carried out on fully saturated soil samples (i.e. dry unit weight, cohesion and friction angle) do not realistically represent actual field conditions. This is because the soil is never placed and compacted in a fully saturated condition during construction. In addition, since the magnitudes of hydraulic conductivity in fine-grained unsaturated soils are extremely low, it is usually unlikely that a significant portion of the slope would ever become saturated even under extreme rainfall conditions. However, in addition to hydraulic conductivity, the degree of saturation in unsaturated soils also depends on the hydraulic gradient which could be significant due to matric suction. Furthermore, saturated soil samples in the laboratory cannot be compacted to the specified relative compaction (e.g. 95% of maximum dry unit weight) to represent field conditions. Therefore, their measured properties would underestimate the corresponding field values. As a result, an embankment design using saturated soil properties will be neither optimized nor realistic. The focus of this paper is on pullout capacity of geotextile reinforcement in marginal soils which is an important design consideration in internal stability of reinforced soil structures. Based on the above discussion, since RSS are typically constructed with the soil compacted at or the vicinity of the optimum (gravimetric) moisture content (OMC ¼ GWCopt), the mechanical properties of the soil for the design of RSS need to be determined in the laboratory at the corresponding GWC values. Therefore, a primary objective of the study is to quantify the magnitude of reduction in the pullout capacity of geotextile reinforcement as a result of loss of matric suction in the unsaturated marginal soil due to wetting. This can lead to excessive deformations and even failure of the reinforced soil structure. However, the influence of the soil GWC on the reinforcement pullout capacity and the resulting factors of safety against failure is not explicitly accounted for in the current design guidelines and provisions. In this study, a moisture reduction factor (MRF), denoted by m(u), is proposed to account for the pullout resistance of geotextile reinforcement in the design of reinforced soil structures with marginal soils. The MRF value is a function of the soil GWC value (and hence of the soil suction), which makes the predicated value of the pullout resistance more accurate and reliable for design purposes. It should be noted that due to the very low hydraulic conductivity of unsaturated marginal soils, measuring the change in the pullout resistance of a significant size soil-geotextile specimen in drained conditions (e.g. conforming to the ASTM D6706 test protocol) is extremely time consuming. Therefore, in this study the marginal soil in the pullout tests described in this paper were placed at prescribed GWC values ranging from OMC-2% to OMCþ2% to determine the corresponding MRF values. Consequently, the MRF values in this study do not exactly represent the reduction in the reinforcement pullout capacity as a result of wetting of a soil mass compacted at an initial GWC and unit weight. However, they provide some quantitative data that could be used to estimate the magnitude of such reduction for design purposes. Furthermore, these results are more directly applicable to determine the expected pullout capacity of the reinforcement in the marginal soil if placed at any GWC value within the range between OMC-2% and OMCþ2%. 2. Theory 2.1. Reinforcement pullout capacity in reinforced soil structures For internal stability, the pullout resistance per unit width (Pr) of the reinforcement in reinforced soil structures is determined using Eq. (1). It is defined as the maximum tensile load required to

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153

Pr ¼ F * as0v Le C

(1)

where: Le: Embedment or adherence length in the resisting zone behind the failure surface C: Reinforcement effective unit perimeter; e.g., C ¼ 2 for strips, grids, and sheets LeC: Total surface area per unit weight of reinforcement in the resistive zone behind the failure surface F* ¼ tandpeak: Pullout resistance (or friction-bearing-interaction) factor dpeak: Equivalent peak friction angle of the soil-geosynthetic interface a: A scale effect correction factor to account for a nonlinear stress reduction over the embedded length of highly extensible reinforcement s0v : Effective vertical stress at the soil-reinforcement interface F*

Pullout tests are typically used to obtain the parameters a and for different combinations of soil and reinforcement materials. Tests are typically performed on samples with a minimum embedded length of 600 mm as recommended in related guidelines (e.g. ASTM D6706). The correction factor a depends on the extensibility and the length of the reinforcement. For extensible sheets (i.e., geotextiles), the recommended value of a is 0.6 (Berg et al., 2009). The parameter F* (especially in reinforcement types such as geogrids and welded wire mesh) includes both passive and frictional resistance components (e.g., Palmeira, 2004; Abu-Farsakh et al., 2005; Berg et al., 2009). It is worth noting that in the case of extensible reinforcement (e.g. geotextiles), a wetter soil results in smaller amount of reinforcement extension before pullout as compared to an otherwise identical soil in a drier condition. In Eq. (1), smaller reinforcement extensibility results in a larger value for a which could point toward a larger Pr value for the case of a wetter soil, which is erroneous. However, the combined term aF* is expected to always decrease with the soil GWC value. Nevertheless, in this paper Eq. (1) is modified in the form:

Pr ¼ F * as0v Le C mðuÞ

Ca0 : Adhesion intercept

sn: Normal stress on the interface ua: Pore air pressure 0 d : Angle of friction between the soil and the structure counterface with respect to (sn  ua) uw: Pore water pressure db: Angle of friction between the soil and the structure counterface with respect to matric suction (ua  uw) In unsaturated soils, Mohr’s circles representing failure conditions correspond to a three-dimensional (3D) failure surface, where the shear stress (s) is the ordinate and the two stress variables (sn  ua) and (ua  uw) are the abscissas. The planar surface formed by these two stress variables is referred to as the extended MohrCoulomb failure envelope (Khoury et al., 2011; Hatami et al., 2013). 3. Laboratory tests 3.1. Materials 3.1.1. Soil classification and properties The soil used for the pullout and interface shear tests in this study was a lean clay found on US Route 62 in Chickasha, OK. The gradation and fines content of the soil were determined using the ASTM D422 (ASTM, 2007) and D1140 (ASTM, 2006) test methods. The test results are given in Fig. 2 and Table 1, which indicate that the soil is classified as CL and A-6, based on the USCS and AASHTO soil classification systems, respectively. The maximum dry unit weight and OMC value of the Chickasha soil from modified Proctor tests were determined as gdmax ¼ 17.3 kN/m3 and OMC ¼ 18%, respectively. A series of direct shear tests (ASTM D3080, 2011) was carried out on the soil at three different GWC values (i.e. OMC-2%, OMC and OMCþ2%) and at a rate of 0.06 mm/min to determine its shear strength parameters (i.e. c0 and B0 ) with the results as shown in Table 1. 3.1.2. Geosynthetic reinforcement A woven polypropylene (PP) geotextile (Mirafi HP370) was used for the pullout and interface shear tests in this study. The mechanical response of the geotextile was found as per the ASTM

(2)

100

by introducing a moisture reduction factor (MRF), m(u), to explicitly account for the influence of the soil GWC value on the soilreinforcement pullout capacity as described in more detail in this paper.

80

60

2.2. Extended Mohr-Coulomb failure envelope According to a theory proposed by Fredlund et al. (1978) the shear strength of an unsaturated soil can be expressed using the following two stress variables: the net normal stress, which is the difference between the total stress and the pore air pressure (sn  ua) and the soil matric suction, which is the difference between the pore air and pore water pressures (ua  uw). Based on the Fredlund et al. (1978) approach, Miller and Hamid (2005) proposed the following equation to determine the shear strength of unsaturated soil-structure interfaces:

ss ¼ Ca0 þ ðsn  ua Þtan d0 þ ðua  uw Þtan db Where:

(3)

40

Percent finer by weight (%)

generate outward sliding of the reinforcement through the reinforced soil mass (Elias et al., 2001; Berg et al., 2009):

141

20

10

1

0.1

0.01

0.001

0 0.0001

Diameter (mm) Fig. 2. Gradation curve of Chickasha soil from sieve analysis and hydrometer tests. The vertical broken line shows the location of the #200 sieve.

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153

Table 1 Summary of Chickasha soil properties. Value Lean clay Liquid limit Plastic limit Plasticity index Specific gravity Gravel (%) Sand (%) Silt (%) Clay (%) Maximum dry unit weight, kN/m3 OMC (%) Cohesion (kPa) at OMC-2%, OMC and OMCþ2% Friction angle ( ) at OMC-2%, OMC and OMCþ2%

38.0 20.0 18.0 2.75 0.0 10.6 49.4 40.0 17.3 18.0 42.6, 29.3, 20.4 29.6, 27.3, 27.1

Mirafi HP370 Ultimate tensile strength (kN/m) Tensile strength at 5% strain (kN/m)

40 20

Note: (1) Maximum dry unit weight and OMC were determined using Modified Proctor compaction effort (ASTM D1557 (2012)); (2) Geotextile properties were determined as per ASTM D4595 (ASTM, 2009) test protocol.

D4595 test protocol (ASTM, 2009) and was compared with the manufacturer’s data (Table 1). 3.2. Suction sensors The soil suction was initially measured using several different methods and sensors including thermal conductivity sensors (Fredlund and Wong, 1989; Fredlund et al., 2000), filter paper tests (ASTM, 2010), PST-55 psychrometer sensors and the WP4 potentiometer. However, based on the range of measured suction values and the accuracy and reparability of the test results, it was concluded that the PST-55 psychrometer and the WP4 potentiometer were the most suitable “in-situ” sensor (i.e. inside the pullout box) and “off-site” equipment to measure the soil suction, respectively. Hence, these instruments are described briefly in the following sections. 3.2.1. PST-55 psychrometers PST-55 is an in-situ psychrometer which can measure soil suction for values up to 5000 kPa. Under vapor equilibrium conditions, the water potential of the PST-55 porous cup is directly related to the vapor pressure of the surrounding air. This means that the soil water potential is determined by measuring the relative humidity of the chamber inside the porous cup (Campbell and Gardner, 1971). PST-55 psychrometers are commonly used in geotechnical research projects. These sensors can lose their factory calibration over time. Therefore, in this study they were calibrated using a 1000 mmol/kg NaCl solution before they were used in the pullout tests. An HR-33T data logger was used to read the water potential of the NaCl solution samples, and an ice chest provided a controlled temperature and moisture environment for the calibration of the sensors. The sensors were submerged in NaCl solutions and kept in the ice chest for 2 h to reach equilibrium (Wescor Inc. 2001). Then, each sensor was connected to the data logger (one at a time) and the voltage representing the water potential of the control NaCl solution was read in microvolts (mV). 3.2.2. WP4 potentiometers The WP4 equipment consists of a sealed block chamber equipped with a sample cup, a mirror, a dew point sensor, a temperature sensor, an infrared thermometer and a fan. The soil sample is placed in the sample cup and brought to vapor equilibrium with the air in the headspace of the sealed block chamber. At equilibrium, the water potential of the air in the chamber is the same as the water potential or suction of the soil sample.

Seventeen (17) 40 mm (diameter) by 6 mm (height) disk-shaped WP4 samples of Chickasha soil were prepared at different GWC values at the same dry unit weight as that in the laboratory pullout tests. The WP4 samples were placed in sealed disposable cups. Before testing each soil sample using WP4, a salt solution of known water potential (i.e. 0.5 molal KCl in H2O) was used to calibrate the WP4 sensor. For each test, the sample was placed inside the WP4 sample cup and was allowed to reach temperature equilibrium with the equipment internal chamber. The magnitude of the soil suction was recorded once the displayed reading stabilized at a constant value. Fig. 3 shows the Soil-Water Characteristic Curve (SWCC) for the Chickasha soil as was obtained from the WP4 tests. Results shown in Fig. 3 indicate that the suction in Chickasha soil varies between 300 kPa and 1200 kPa for the range of GWC values between OMC-2% and OMCþ2%. This range of soil suction is consistent with the values reported in the literature for lean clay (e.g. Cardoso et al., 2007; Nam et al., 2009). 3.3. Large-scale pullout tests A series of large-scale pullout tests were carried out on the woven geotextile in Chickasha soil. The tests were carried out at three different GWC values (Table 2). The differences in the magnitude of geotextile pullout resistance among these cases were used to determine a moisture reduction factor (MRF), denoted by m(u) in Eq. (2), to account for the loss of reinforcement pullout resistance due to an increase in the soil GWC value. 3.3.1. Test equipment The nominal dimensions of the large-scale pullout test box used in this study (Fig. 4) are 1800 mm (L)  900 mm (W)  750 mm (H). The size of the box and its basic components, including metal sleeves at the front end exceed the minimum requirements of the ASTM D6706 test standard (ASTM, 2013). The boundary effects in the test were further minimized by lining the sidewalls of the test box with plastic sheets to reduce sidewall friction, and placing Styrofoam panels in contact with the soil inside the box to provide added compressibility at the front boundary. A surcharge assembly including an airbag and reaction beams across the top of the soil surface was used to apply overburden pressures up to 50 kPa on the soil-reinforcement interface. The pullout load on the reinforcement specimen was applied using a 90 kN high-precision servocontrolled hydraulic actuator. In the tests carried out in this study, only one half of the box length (i.e. 900 mm) was used.

30 25 20 GWC (%)

142

15 10 5 0 100

1000 Soil suction (kPa)

10000

Fig. 3. Soil water characteristic curve for Chickasha soil using WP4 potentiometer. Dashed lines show GWC at OMC-2% and OMCþ2%.

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153 Table 2 Large-scale pullout test parameters. Test information Soil Geosynthetic reinforcement Overburden pressure, kPa Gravitational water content (GWC)

Chickasha soil Mirafi HP370, woven PP 10, 20, 50 OMC-2%, OMC, OMCþ2%

3.3.2. Instrumentation Different instruments were used in the pullout tests to measure the movement of geotextile reinforcement and the matric suction in the soil. Deformation of the geotextile reinforcement was measured using four (4) wire-line extensometers attached to

143

different locations along its length (Fig. 5a). A Geokon Earth Pressure Cell (EPC) was used to verify the magnitude of the overburden pressure on the soil-geotextile interface from the airbag that was placed on the top of the soil (Fig. 5b). PST-55 psychrometers were placed 25 mm above and below the soil-geotextile interface to measure and monitor the soil suction near the soil-reinforcement interface. The locations of WP4 samples and those of the in-situ PST-55 psychrometers are shown in Fig. 6 using white and black circles, respectively. 3.3.3. Test procedure The soil was air dried and its larger clumps were broken into smaller pieces. It was then grinded into smaller pieces and passed

Air tube to fill up air bag

Plastic sheets to minimize sidewall friction

Styrofoam panels to fill space above soil specimen 750 mm Geotextile reinforcement attached to actuator

1800 mm 900 mm

Vertical columns to limit horizontal deformation of test box Fig. 4. One of the pullout test boxes at the OU Geosynthetics Laboratory.

Pullout Direction

(a)

(b)

Fig. 5. (a) Wire-line extensometers attached to the geotextile reinforcement. L is distance from the front end of the geotextile; (b) Earth pressure cell on the top of the soil in the pullout test box.

144

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through a #4 sieve using a soil processor. Afterward, the soil was mixed with water to reach the desired GWC value for each test. The wet soil was stored in sealed buckets for at least 24 h to reach moisture equilibrium. The soil GWC value in each bucket was measured by testing one soil sample using the oven drying method. The above procedure was repeated for every pullout test. After the soil was processed and was ready to be placed in the test box, the pullout box was lined with plastic sheets to preserve the soil water content and to minimize the friction between the soil and the sidewalls during each test. Next, the soil was placed and compacted in the test box in nine 50 mm lifts. The thickness of the soil layers below and above the geotextile reinforcement was 230 mm which exceeded the minimum depth of 150 mm according to the ASTM D6706 test protocol (ASTM, 2010a,b). The soil was compacted to 95% of its maximum dry unit weight (i.e. gd ¼ 16.4 kN/m3). The instrumented geotextile was placed at the mid-height of the box. The pullout box containing compacted soil at its target GWC value was sealed with plastic sheets on the top. The soil was left for at least 24 h until the psychrometer sensors reached equilibrium with their surrounding soil and for three to four additional days until the soil GWC inside the pullout test box stabilized. In all pullout tests, rectangular Styrofoam blocks with dimensions 900 mm (L), 228 mm (H) and 140 mm (T) were placed in front of the soil specimen above and below the 200 mm-wide metal sleeves, which helped further minimize the influence of the front boundary condition on the soil-geotextile interface. The pullout force was applied on the geotextile reinforcement at a target displacement rate of 1 mm/min according to the ASTM D6706 test protocol. 3.4. Results of large-scale tests and discussion 3.4.1. Water content and suction Figs. 7 and 8 show distributions of the soil GWC and suction in each layer for the large-scale pullout tests carried out at different GWC and subjected to 50 kPa overburden pressure. The mean and Coefficient of Variation (COV) values for these parameters were calculated for the fifth layer (lift) in the pullout box (i.e. for the soil layer in contact with the geotextile reinforcement) to examine the proximity of their as-placed and target values. Table 3 shows the mean and COV values for the GWC and suction in the fifth layer in large-scale pullout tests. The accuracy of the soil suction values from the PST-55 psychrometers was also examined by comparing them with the readings from the WP4 potentiometer as shown in

Fig. 6. Schematic diagram of the large-scale pullout test box setup (not to scale). Notes: (1) Black and white circles represent the locations of PST-55 sensors and soil samples for the WP4 sensor, respectively; (2) The distance between the sensors and the interface is 25 mm; (3) The sleeves above and below the geotextile layer are 200 mm wide.

Table 4. The GWC and suction COV values for all test cases as given in Figs. 7 and 8 and Table 3 are overall reasonable and indicate that the soil moisture condition was fairly uniform and consistent throughout the large-scale test models. 3.4.2. Reinforcement strain and interface strength Fig. 9 shows the strain distributions over the length of geotextile reinforcement at maximum pullout force based on the extensometer results for the test case under 50 kPa overburden pressure. The strain near the front end of the geotextile reinforcement was calculated by subtracting the displacements measured at the location of Extensometer 1 from those measured at the front end of the geotextile exiting the soil. The displacements at the front end were determined by subtracting the calculated elongation of the inair portion of the geotextile specimen from the actuator displacement. Results in Fig. 9 indicate that strains in the geotextile reinforcement are greater at higher overburden pressures and lower GWC values (i.e. higher soil suction). Fig. 10 shows the pullout test data and results of interface shear strength (s) for the Chickasha soil for different magnitudes of GWC and overburden pressure. The s values were calculated by dividing the pullout force for each case by the in-soil area of the geotextile specimen (i.e. two times the geotextile area). The target GWC values in the pullout tests include OMC-2% (16%), OMC (18%) and OMCþ2% (20%) (see Table 1). In Fig. 8a-c, the measured pullout force is plotted as a function of the actuator displacement. Results shown in Fig. 10 quantify the increase in the reinforcement pullout resistance in Chickasha soil with overburden pressure for a given GWC value. It should be noted that for the test at OMC-2% subjected to 50 kPa overburden pressure (Fig. 10a) it was found that the geotextile had been ruptured before pullout. Therefore, the pullout force at failure was estimated using the trends in the corresponding test data at OMC and OMCþ2%. As expected, increasing suction led to a higher maximum reinforcement pullout resistance in otherwise identical test specimens (Fig. 10d and e). Results shown in Fig. 10d indicate that apparent adhesion increases at lower GWC due to higher suction. This observation is consistent with those reported by Khoury et al. (2011) from suction-controlled interface testing of fine-grained soil specimens. Results shown in Fig. 10a-d represent the frontal planes of extended Mohr-Coulomb failure envelopes for the soil-geotextile interface at different GWC and suction values. These failure envelopes can be considered to be practically linear for all GWC cases examined. The interface strength results, i.e. the values for the slope (tan d0 ) and the intercept (ca) of the failure envelopes on these frontal planes are summarized in Table 5. Abu-Farsakh et al. (2007) studied the effect of the GWC on the interaction between three cohesive soils and a woven geotextile reinforcement material. They found that an increase in the molding GWC of the soil from 24% to 33% caused 43% reduction in the interface shear resistance. The data summarized in Table 5 are overall consistent with Abu-Farsakh et al.’s observations. For instance, the pullout resistance (Pr) at OMCþ2% is between 17% and 35% lower than the corresponding value at OMC-2% depending on the overburden pressure. A smaller confining pressure resulted in a greater reduction in pullout resistance for a given increase in the soil GWC value. Results in Fig. 10e show the failure envelopes of the threedimensional extended Mohr-Coulomb failure surface on the lateral plane for the soil-geotextile interface as a function of the soil suction. The line intercept and slope represent the effective adhesion at zero overburden pressure (sn ¼ 0 kPa) and interface friction angle with respect to suction (db), respectively. The data shown in Fig. 10e indicate that the interface friction angle with respect to suction for the Chickasha soil-geotextile tested is negligible (it is less than 1 ; note the significantly different scales of the horizontal

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153

and vertical axes in the figure). These results indicate that as the overburden pressure increases, interface adhesion and consequently, the interface shear strength increases. The extended MohrCoulomb envelope in Fig. 11 shows the variation of the interface shear strength (s) with the values of soil suction and overburden

24

20

GWC (%)

145

16 1600

12

8

Mean GWC per layer 4 1

2

3

4

5

6

7

8

9

10

Soil lift number in pullout box

Total suction (kPa)

1200

GWC for each bucket

800

OMC-2%

400

(a)

v

24

Soil suction distrinution per layer

: 50 kPa

Mean soil suction per layer 125 kPa

COV: 3.4 0 1

2

3

4

5

6

7

8

9

10

20 Soil lift number in pullout box

GWC (%)

(a) 16

1600 OMC

Soil suction distribution per layer

σv : 50 kPa

12

Mean soil suction per layer

18% GWC for each bucket

8

v:

50 kPa Mean GWC per layer

COV: 1.5% 4 1

2

3

4 5 6 7 Soil lift number in pullout box

8

9

10

Total suction (kPa)

1200

COV: 8.2%, Ψ = 590 kPa

800

400

(b) 24 0 1

2

3

20

4

5

6

7

8

9

10

Soil lift number in pullout box

OMC+2% v

12

1200

20.1% 8

v:

GWC for each bucket

50 kPa

Mean GWC per layer

COV: 1.7%

4 1

2

3

4

5

6

7

8

9

10

Soil lift number in pullout box (c) Fig. 7. Distributions of the soil GWC with depth in the pullout box for different pullout test cases. Notes: (1) One soil sample was taken from each bucket to test its GWC value; (2) The number of soil samples from each soil lift in the pullout box is given in Table 5 (caption); (3) The horizontal line indicates the target GWC for each test case; (4) The vertical dashed line shows the location of the soil-geotextile interfaces; (5) The mean and COV values reported in the legends are calculated for the fifth layer (i.e. soilgeotextile interface) data only.

Total suction (kPa)

GWC (%)

(b) 1600

16

Soil suction distribution per layer

: 50 kPa Mean soil suction per layer

COV: 11

298 kPa

800

400

0 1

2

3

4 5 6 7 Soil lift number in pullout box

8

9

10

(c) Fig. 8. Distributions of the soil suction with depth in the pullout box from WP4 at different GWC. The number of soil samples from each soil lift in the pullout box is reported in Table 5.

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Table 3 Mean and COV values for the GWC and suction in the fifth layer (in contact with geotextile) in large-scale pullout tests. Target u (%)

sn (kPa) Mean J (kPa) COV(J) (%) Mean u (%) COV(u) (%)

16 (OMC-2%)

10 20 50

1236 1196 1125

7.3 5.9 3.4

15.7 15.8 16.0

1.5 1.1 0.7

18 (OMC)

10 20 50

513 570 590

5.7 6.7 8.2

18.5 18.1 18.0

0.9 1.1 1.5

10 20 50

304 352 298

11.2 9.9 11.0

20.1 19.6 20.1

1.8 1.6 1.7

20 (OMCþ2%)

multi-scale laboratory testing approach. It is worth noting that both the soil particle size and the asperities of the geotextile reinforcement are orders of magnitude smaller than the dimensions of the small-scale text box. Therefore, the test results are not believed to be negatively impacted by scale effects. However, proper boundary

30 OMC-2%, 10 kPa OMC, 10 kPa

25

OMC+2%, 10 kPa

Geotextile reinforcement

3.5. Small-scale tests In addition to large-scale pullout tests, a series of small-scale pullout and interface shear tests were performed on the same soil that was used in the large-scale pullout tests. In addition, these tests were carried out at the same soil GWC, unit weight (i.e. 95% of maximum dry unit weight from modified Proctor tests) and overburden pressure magnitudes as those in the large-scale pullout tests (Table 7). One main advantage of small-scale pullout tests is that once they are calibrated against the large-scale tests, they can be carried out at significantly larger numbers to test a large variety of soil types, GWC values and overburden pressures. This will help to develop a better understanding of the influence of the soil GWC and matric suction on marginal soil-geotextile interfaces using a

Locations where potentiometers were attached to the geotextile

15 10 5 0 0

100 200 300 400 500 Distance on the geotextile from front end of the soil (mm)

30 OMC-2%, 20 kPa OMC, 20 kPa

25

OMC+2%, 20 kPa 20 15 10 5 0 0

100 200 300 400 500 Distance on the geotextile from front end of the soil (mm)

30 OMC-2%, 50 kPa OMC, 50 kPa

sn (kPa)

Mean J (kPa) WP4a

Mean J (kPa) PST-55b

16 (OM-2%)

10 20 50

1243 1200 1135

921 906 910

18 (OMC)

10 20 50

520 570 580

488 520 543

20 (OMCþ2%)

10 20 50

311 360 282

303 335 310

a Mean values were calculated using four undisturbed samples from the fifth layer (in contact with geotextile) for each pullout test as shown in Fig. 6. b Mean values were determined using three PST-55 psychrometers placed in the fifth soil layer (Fig. 6).

OMC+2%, 50 kPa 20

Strain (%)

Target u (%)

600

(b)

25 Table 4 Comparison of suction values in Chickasha soil as measured using psychrometers (in-situ) and WP4 (offsite equipment).

600

(a)

Strain (%)

pressure at the soil-reinforcement interface. Taken together, the results based on the description of soil shear strength using two stress state variables (i.e. soil suction and net normal stress) as presented in Figs. 10 and 11 and Table 5 are in good agreement with those reported by Hatami et al. (2010a) and Khoury et al. (2011) on other marginal soils. Table 6 shows aF* (Eq. (1)) values calculated from all large-scale pullout tests in Chickasha soil. Example calculations for a according to the FHWA guidelines (Berg et al., 2009) are shown in Fig. 12. Based on the intersection of the horizontal asymptote with the yaxis, the design value for a from pullout tests in Chickasha soil was found to be 0.5 (Fig. 12e). This value of a indicates a fairly extensible geotextile material and a linear strain distribution along its length. It is also comparable to the value reported by Hatami et al. (2010a) for the same geotextile material tested in Minco silt (i.e. a ¼ 0.59) and the value a ¼ 0.6 recommended by FHWA for geotextiles (Berg et al., 2009). F* values were calculated using Eq. (1).

Strain (%)

20

15 10 5 0 0

100 200 300 400 500 Distance on the geotextile from front end of the soil (mm)

600

(c) Fig. 9. Axial strain distributions in geotextile reinforcement subjected to pullout load from large-scale pullout tests on Chickasha soil at different GWC values.

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153

147

ll

l

l

l

(c)

(b)

(d)

ll

(a)

(e) Fig. 10. Pullout test data and interface strength results from large-scale pullout tests for Chickasha soil at different GWC values: (a)-(c) Load-displacement data; (d) Failure envelopes for the soil-geotextile interface on frontal plane; (e) Failure envelopes for soil-geotextile interface on lateral plane. Note: in (a), dashed line indicates the estimated pullout failure.

conditions are required to help calibrate these test results against those from large-scale pullout tests. The small-scale pullout and interface shear tests were carried out using a direct shear testing (DST) machine shown in Fig. 13. The soil in both tests was placed in a 60 mm  60 mm square test cell of

the DST machine. Two rectangular blocks of Styrofoam with dimensions 60 mm (L), 12 mm (H) and 9 mm (T) were used in the small-scale pullout tests in front of the soil specimen to provide a compressible boundary condition similar to that in the large-scale pullout box. The Styrofoam blocks were placed in the upper and

Table 5 Interface strength properties from large-scale pullout tests in Chickasha soil. 0

Target u (%)

sn (kPa)

Mean u (%)a

Mean J (kPa)b

Pr (kN/m)

smax (kPa)

d ( )

Ca (kPa)

16 (OMC-2%)

10 20 50

16.0 16.0 16.0

1153 1151 1135

29.6 34.8 45.2

24.3 28.5 37.1

17.3

21.6

18 (OMC)

10 20 50

18.3 18.2 18.1

550 566 576

24.8 29.7 38.7

20.3 24.4 31.7

15.4

18.1

20 (OMCþ2%)

10 20 50

20.3 20.0 20.2

286 312 290

19.1 28.8 33.8

15.7 23.6 27.7

14.7

15.3

a b

Mean values were calculated using 45 GWC samples for each pullout test (5 samples from each of the nine 2-inch soil lifts). Mean values were determined from SWCC for Chickasha soil (Fig. 3) based on GWC values determined for each test (i.e. 45 data points).

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Fig. 11. Extended Mohr-Coulomb envelope from large-scale pullout tests.

lower halves of the test cell in front of the soil. A 20 mm (W)  40 mm (L) geotextile specimen was used in each pullout test. The linear scale factor between the small-scale and large-scale pullout tests was 1:15. In the small-scale pullout tests, the geotextile specimen was pulled out of a fixed test cell filled with Chickasha soil at a speed of 0.06 mm/min (i.e. 1/15 of 1 mm/min nominal rate at large scale). In the interface shear tests, the lower box of the DST machine was pushed horizontally at a speed of 0.06 mm/min to apply a shear load on the soil-geotextile interface. These two tests are described in more detail in the following sections. 3.5.1. Small-scale pullout tests The soil in the small-scale pullout tests was prepared using the same process as was followed for the large-scale tests: The clay was first processed, then passed through a #4 sieve (i.e. 4.75 mm aperture size) and mixed with water to reach the target GWC value which was measured using the oven drying method preceding and following each test. The soil was placed in the bottom half of the test cell in four lifts at the target GWC value and each lift was compacted to the final thickness of 3 mm. The geotextile specimen was attached to a custom-made clamp mounted on the test box and was embedded 40 mm inside the test cell. A U-shaped metal spacer (open on the front side) was used to maintain a gap within the pullout slot to prevent any frictional contacts within the test cell frame during the pullout process. The top half of the box was filled with 4 more layers of the soil, each compacted to 3 mm thickness.

Table 6 Calculated values of aF* from large-scale pullout tests in Chickasha soil. Target u (%)

sn (kPa)

Pr (kN/m)

smax (kPa)

aF*

16 (OMC-2%)

10 20 50

29.6 34.8 45.2

24.3 28.5 37.1

2.43 1.43 0.74

18 (OMC)

10 20 50

24.8 29.7 38.7

20.3 24.4 31.7

2.04 1.22 0.64

20 (OMCþ2%)

10 20 50

19.1 28.8 33.8

15.7 23.6 27.7

1.57 1.18 0.56

3.5.2. Results Fig. 14 shows the plots of pullout force versus actuator displacement for the small-scale pullout tests in Chickasha soil. Soil-geotextile interface strength properties obtained from the small-scale tests are summarized in Table 8. The pullout test results given in Table 8 and Fig. 14 show a clear influence of the soil overburden pressure and GWC on the soilgeotextile interface strength and pullout resistance. It is observed that the pullout force increases with overburden pressure. The interface adhesion contributing to the geotextile pullout resistance decreases by 31% as the soil GWC increases from OMC-2% to OMCþ2%. The interface friction angle also decreases by 40% from OMC-2% to OMCþ2%. These results are consistent with those obtained by the authors in a previous study on Minco silt (Hatami et al., 2010a,b). Results in Figs. 14d and 8d indicate that the interface adhesion from both small-scale and large-scale pullout tests depends on the soil GWC and it is consistently larger for greater soil suction values (i.e. lower GWC). These results also indicate that the magnitudes of soil-geotextile interface adhesion from small-scale pullout tests are greater than those from the corresponding largescale tests. This could be attributed to the smaller size and greater boundary effects in the small-scale tests. Consequently, a calibration (or scale) factor needs to be determined and applied to the small-scale test results before they can be used for practical applications. The data in Fig. 14e indicate that the interface shear strength increases with overburden pressure as a result of increase in interface adhesion ðCa0 Þ. The results shown in Fig. 14e indicate that the interface friction angle with respect to suction (db) for the Chickasha soil-geotextile tested is less than 2 . 3.5.3. Interface shear tests A series of interface shear tests was carried out on the woven geotextile with Chickasha soil at different GWC values to determine the strength properties of the Chickasha soil-geotextile interface. A geotextile specimen was attached to a 60 mm  60 mm aluminum panel and was placed on the top of a stack of aluminum panels in the bottom half of the test cell in the DST machine. The top half of the test cell was filled with four 3 mm-thick compacted layers of Chickasha soil, similar to the small-scale pullout tests. 3.5.4. Results Fig. 15 shows the Mohr-Coulomb envelopes from interface shear tests at different GWC values. The results show that the soil-geotextile interface strength increases consistently with the overburden pressure and with the soil matric suction. According to Fig. 15a, the interface friction angle was found to decrease by 15% from OMC-2% to OMCþ2%. Results in Fig. 15b indicate that the interface friction angle with respect to matric suction on the lateral plane is less than 1, which is consistent with the data from largescale pullout tests (Fig. 10e). Fig. 16 shows the extended MohrCoulomb envelopes from small-scale pullout and interface shear tests. The plots of Mohr-Coulomb envelopes in Fig. 16 show that the adhesion values calculated from pullout tests are greater which could be attributed to this fact that, in pullout tests, as opposed to the interface shear tests, the geotextile is stretched during the test. This could result in the enlargement of the geotextile openings which, in turn, could allow the fine-grained soil to penetrate into the plane of the geotextile. Similar to geogrids but at a smaller scale, the soil within the openings of the geotextile subjected to overburden pressure could exhibit some passive resistance against the pullout force which could be responsible for the larger adhesion intercept that is observed for the pullout test results as compared to the interface shear data. Table 9 and Fig. 17 summarize the data from all laboratory tests carried out in this study.

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l

l

l

(c)

ll

ll

(a)

l

(b)

(d)

l l

(e) Fig. 12. Calculation of pullout parameters for Mirafi HP370 geotextile reinforcement from large-scale pullout test data in Chickasha soil at OMCþ2% subjected to 50 kPa overburden pressure: (a) Pullout force versus actuator displacement; (b) Pullout force versus extensometer displacement; (c)-(e) Procedure to determine F* and a using large-scale pullout tests; Note: in (b), EX1 and EX4 are near the front end and the tail end, respectively. In (c), solid and dashed lines indicate actual and interpolated data, respectively.

4. Moisture reduction factor, m(u) Fig. 18 shows the variations of m(u) for the Chickasha soilwoven geotextile interface as a function of the soil GWC at

Table 7 Small-scale pullout and interface test parameters. Test information

Chickasha soil

Type of small-scale test Geosynthetic reinforcement Soil specimen dimensions

Pullout, Interface shear Mirafi HP370, woven PP 60 mm (W)  60 mm (L)  24 mm (H), & 60 mm (W)  60 mm (L)  12 mm (H) 20.3 mm (W)  40.6 mm (L) & 60 mm (W)  60 mm (L) 10, 20, 50 OMC-2%, OMC, OMCþ2%

Geosynthetic reinforcement dimensions Overburden pressure, kPa Gravitational water content (GWC)

different overburden pressures from all three categories of tests carried out in this study. In the calculation of m(u), the interface strength at u ¼ OMC-2% is taken as the reference value (Hatami et al., 2010a,b). Results shown in Fig. 18 indicate that construction of reinforced soil slopes and embankments on the wet side of OMC or wetting of the soil-geotextile interface during construction or service life of the reinforced soil structure (as compared to e.g., the case of OMC2%) could result in considerably lower pullout resistance of the geotextile reinforcement. The calculated amounts of reduction in interface strength from OMC-2% to OMCþ2% as obtained from the large-scale and small-scale test data are between 17% and 42% depending on the test cases and overburden pressures applied on the specimens. Results shown in Fig. 18 indicate that the variation of m(u) with the soil GWC could be approximated as linear for practical purposes for the range of GWC values examined in this study.

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ll

ll

Fig. 13. Small-scale pullout tests in Chickasha soil using a DST machine.

l

l

(c)

ll

(a)

l

(b)

(d)

(e) Fig. 14. Pullout test data and interface strength results from small-scale tests for Chickasha soil and comparison of failure envelopes for soil-geotextile interface at different GWC values: (a)-(c) Load-displacement data; (d) Failure envelopes for soil-geotextile interface on frontal plane; (e) Failure envelopes for soil-geotextile interface on lateral plane. Note: Suction values were calculated from the SWCC.

D. Esmaili et al. / Geotextiles and Geomembranes 42 (2014) 139e153

151

Table 8 Interface strength properties from small-scale pullout tests. Target u (%)

sn ( kPa)

u (%)

J (kPa)

smax ( kPa)

d0 (  )

Ca (kPa)

16 (OMC-2%)

10 20 50

16.3 15.9 16.0

1036 1164 1131

56.5 76.3 83.8

30.2

56.7

18 (OMC)

10 20 50

17.8 17.9 18.3

633 613 538

46.8 53.9 65.2

23.8

43.5

20 (OMCþ2%)

10 20 50

20.0 19.9 19.8

310 320 331

40.6 47.5 54.7

18.1

38.9

Note: Suction values were calculated from the SWCC.

5. Conclusions The primary objective of this study was to develop a moisture reduction factor (MRF), denoted as m(u), for the pullout resistance of geotextile reinforcement for the design of reinforced soil structures with marginal soils. Based on the results of this study, the current FHWA design equation for pullout capacity of

Fig. 16. Extended Mohr-Coulomb envelope from small-scale pullout and interface shear tests.

Table 9 Summary of the results from all laboratory tests performed in this study.

100

Type of test

OMC-2% OMC

Shear strength (kPa)

80 δ' = 20º ca = 27 kPa

δ' = 21º ca = 34.8 kPa

60

OMC+2%

Large-scale pullout

Small-scale pullout

Interface shear

Target u (%)

sn (kPa)

d0 ( )

Ca (kPa)

d0 (  )

Ca (kPa)

d0 (  )

Ca (kPa)

16 (OMC-2%)

10 20 50

17.3

21.6

30.2

56.7

21.0

34.8

18 (OMC)

10 20 50

15.4

18.1

23.8

43.5

20.0

27.0

20 (OMCþ2%)

10 20 50

14.7

15.3

18.1

38.9

18.0

21.2

40

20 δ' = 18º ca = 21.2 kPa 0 0

20

40

60

Normal stress (kPa)

(a)

80 60

60

δb = 0.9º ca ' = 29.0 kPa

Small-scale pullout tests

40 10 kPa 20

20 kPa

δb = 0.8º ca '= 18.6 kPa

0 0

500

50 kPa 1000

40 Interface adhesion

1500

30 40

20

20

Small-scale interface shear tests 10 Large-scale pullout tests

Suction (kPa)

(b)

0

0 12

Fig. 15. Mohr-Coulomb envelopes for Chickasha soil-geotextile interface from interface shear tests: (a) Envelopes on the frontal plane; (b) Envelopes on the lateral plane. Note: Suction values were calculated from the SWCC.

Interface friction angle Interface friction angle (°)

δb = 1º ca ' = 32.3 kPa

Interface adhesion (kPa)

Shear strength (kPa)

100

geotextile reinforcement was modified to explicitly account for the influence of the marginal soil gravitational water content (GWC) on the soil-reinforcement interface strength. The m(u) values were determined through a series of multi-scale pullout and interface tests in a silty clay in the laboratory. The test results indicated that the change in the soil suction as a result of variation in its GWC value could have a significant

16

20

24

Normal stress (kPa) Fig. 17. Comparison of large-scale and small-scale pullout and interface test data.

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other combinations of marginal soil and geotextile reinforcement for field applications. It should be noted that the behavior of a marginal soil which is initially placed and compacted at OMC-2% (with a flocculated structure) and is wetted to OMCþ2% is different from the same marginal soil placed and compacted at OMCþ2% (with a dispersed structure; e.g. Lambe, 1958; Fredlund and Rahardo, 1993). Consequently, the values of m(u) for a soil initially compacted at a lower GWC and then wetted to a larger target GWC value are expected to be somewhat different from those given in Fig. 18. Nevertheless, the broader conclusions of the study and their implications to the design of SRW walls with marginal soils are believed to remain valid.

1.2 μ10 (ω) = -0.0885ω + 2.42 μ20 (ω) = -0.0430ω + 1.67 μ20 (ω) = -0.0633ω + 2.01

μ (ω)

0.9

0.6

Large-scale pullout tests, 10 kPa

0.3

Large-scale pullout tests, 20 kPa Large-scale pullout tests, 50 kPa

0.0 14

16

18 20 ωcompaction (%)

22

24

The authors would like to acknowledge the funding and support from the Oklahoma Department of Transportation (ODOT), the Oklahoma Transportation Center (OkTC) and TenCate Geosynthetics for the study reported in this paper. Contributions of Mr. Michael Schmitz at the Fears Structural Laboratory and undergraduate students: Brandi Dittrich, John Tucker, Kyle Olson, Carlos Chang, Thai Dinh and Jesse Berdis in this project are also acknowledged.

(a)

1.2 μ10 (ω) = -0.0704ω + 2.12 μ20 (ω) = -0.0944ω + 2.47 μ50 (ω) = -0.0868ω + 2.37

μ (ω)

0.9

References

0.6

Small-scale pullout tests, 10 kPa

0.3

Small-scale pullout tests, 20 kPa Small-scale pullout tests, 50 kPa

0.0 14

16

18 20 ωcompaction (%)

22

24

(b)

1.2 μ10 (ω) = -0.105ω + 2.68 μ20 (ω) = -0.0758ω + 2.20 μ50 (ω) = -0.0818ω + 2.31

0.9

μ (ω)

Acknowledgments

0.6 Small-scale interface tests, 10 kPa

0.3

Small-scale interface tests, 20 kPa Small-scale interface tests, 50 kPa

0.0 14

16

18 20 ωcompaction (%) (c)

22

24

Fig. 18. Moisture reduction factor for the woven geotextile in Chickasha soil: (a) Largescale pullout tests; (b) Small-scale pullout tests; (c) Small-scale interface shear tests.

influence on the soil-geotextile interface strength. The shear strength of the soil-geotextile interface at OMCþ2% was between 17% and 42% lower than that of an otherwise identical interface at OMC-2%. The results reported in this paper were obtained using one combination of marginal soil and woven geotextile only. Similar tests and analyses are underway to develop MRF values for

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