Shiping Yang,1 Robert A. Lohnes,2 and Bruce H. Kjartanson3

Mechanical Properties of Shredded Tires

REFERENCE: Yang, S., Lohnes, R. A., and Kjartanson, B. H., “Mechanical Properties of Shredded Tires,” Geotechnical Testing Journal, GTJODJ, Vol. 25, No. 1, March 2002, pp. 44–52. ABSTRACT: The use of scrap tires as construction materials in civil engineering applications is one of the most promising ways of recycling this troublesome waste material. Design of scrap tire structures, however, requires data on engineering characteristics of tire-derived materials. Confined compression, direct shear, and triaxial tests were carried out to evaluate the mechanical characteristics of tire chips approximately 2 to 10 mm in size. These test results were synthesized with data from previous shredded tire studies to generate empirical relationships between normal stress and direct shear strength and between confining pressure and initial tangent modulus from triaxial testing. It was found that the shear strength of shredded tires is independent of the particle size of the material, and the strength envelope is a power function for normal stresses from 0 to 90 kPa. The initial tangent modulus relates to confining pressure through a quadratic equation, and the lateral strain ratio is independent of confining stress. KEYWORDS: scrap tires, shredded tires, tire chips, tire shreds, direct shear test, triaxial test, Young’s modulus, friction angle, cohesion, Mohr-Coulomb envelope

Shredded scrap tires have been used as construction materials in civil engineering applications including retaining wall backfill (Humphrey 1993), road embankments (Bosscher et al. 1993), and subsurface drainage systems (Kjartanson et al. 1998). Structures utilizing shredded scrap tires, however, should be designed to minimize the potential for internal heating and combustion. This is particularly important when the tire shreds are not mixed with soil. The guidelines presented in ASTM Practice for Use of Scrap Tires in Civil Engineering Applications D 6270 (1998) should be followed to minimize this risk. The design of the above-mentioned structures requires engineering properties of the shredded tire material such as shear strength, compressibility, and stress-strain response. The properties of materials ranging in size from less than 10 to 1400 mm are examined. The terminology of ASTM D 6270 is generally followed; however, particles less than about 12 mm in size, termed granulated or ground rubber, are grouped with the tire chips (particles from 12 to 50 mm in size as defined in ASTM D 6270), and particles greater than 50 mm in size are grouped with tire shreds (particles from 50 to 305 mm as defined in ASTM D 6270). 1 Project engineer, TEAM Services, 333 SW 9th Street, Suite H, Des Moines, IA 50309. 2 University professor, Department of Civil and Construction Engineering, Iowa State University, Ames, IA 500113232. 3 Associate professor, Department of Civil and Construction Engineering, Iowa State University, Ames, IA 50014.

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Humphrey et al. (1993), Foose et al. (1996), and Gebhardt (1997) conducted direct shear tests on tire chips and shreds ranging in size from 38 to 1400 mm. Using peak shear stress, or stress at a horizontal displacement equal to 10 or 9% of the length of the shear box if no peak stress is observed, as a failure criterion, the shear strength parameters for Mohr-Coulomb failure envelopes were calculated. Friction angles from these studies ranged from 19 to 38° with cohesion of 0 to 11.5 kPa at normal stresses between 0 and 83 kPa as shown in Table 1. Previous studies interpreted the results according to accepted soil mechanics principles; however, Gebhardt (1997) suggested a power function to describe the relationship between direct shear strength () and normal stress ():   1.40.79

(1)

Bressette (1984), Ahmed (1993), Benda (1995), Masad et al. (1996), Wu et al. (1997), and Lee et al. (1999) conducted triaxial tests on tire chips from 2 to 51 mm in size. Tests were conducted in compression-loading except Wu et al. (1997), who, in addition, conducted compression-unloading tests where the confining pressure 3 was reduced in increments from the initial consolidation pressure while simultaneously increasing the vertical load to keep 1 constant. A linear stress-strain response up to 30% strain was observed from all compression loading tests at confining pressures between 35 and 350 kPa. Because strain softening was not observed, deviator stresses at 10 or 20% axial strain were selected as the shear strength for determining Mohr-Coulomb parameters. Friction angles range from 6 to 57°, and cohesions vary from 0 to 82 kPa as shown in Table 2. The initial tangent modulus of the stress-strain curves, analogous to Young’s modulus, ranges between 300 and 2500 kPa with higher values at higher confining stresses. The objective of this study was to conduct direct shear tests, triaxial tests, and confined compression tests on tire chips and synthesize these results with the existing data in order to find a unified method to quantify the shear strength and stress-strain response of shredded tire materials. Materials Tested The tire chips used in this study were “6 plus tire crumb” provided by EnTire Recycling, Inc. of Nebraska City, Nebraska. This material contains no steel wires but has a small amount of nylon fibers. The chips are poorly graded with particle sizes ranging from 2 to 10 mm. The specific gravity of the tire chips, obtained with a helium pycnometer, averaged 1.15 for three tests. For comparison, the apparent specific gravity of various tire shreds and tire chips reported in ASTM D6270 varies from 1.02 to 1.27 with an average of 1.15.

Copyright © 2002 by ASTM International, 100 Barr Harbor Drive, PO Box C700, West Conshohocken, PA 19428-2959.

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TABLE 1—Shear strength of tire chips and shreds from direct shear testing.

Author Humphrey et al. (1993) Foose et al. (1996) Gebhardt (1997) This study

Maximum Size, mm

Unit Weight, kN/m3

Normal Stress, kPa

Cohesion Intercept, kPa

Friction Angle, 

Criterion of Failure Stress

51 76 38 50, 100, 150 1400 10

6.30 6.08 6.06 NA NA 5.73

17–68 17–63 17–62 1–76 5.5–28 0–83

7.7 11.5 8.6 3 0 0

21 19 25 30 38 32

Peak or at 10% disp.a Peak or at 9% disp.a 10% disp. 10% disp.

a The failure was considered to be the peak shear stress or, if no peak was reached, the shear stress at a horizontal displacement equal to 10% (or 9%) of the length of shear box was taken.

TABLE 2—Shear strength of tire chips and shreds from triaxial testing.

Sample Preparation and Test Methods All the specimens were weighed and divided into two or three portions of equal weight. The tire chips were then poured into the triaxial mold in three lifts or into the direct shear box and confined compression test chamber in two lifts. Each lift of tire chips was rodded vertically 40 times with a spatula. The initial unit weight and void ratio of the samples in this study were 5.73 kN/m3 (0.585 g/cm3 or 36.5 pcf) and 0.98, respectively. Two confined compression tests were performed in an oedometer with a sample chamber of 63.5 mm inside diameter and 25.4 mm height. The height of the initial sample is less than ten times the maximum particle diameter as required by the ASTM Test

Method for One-Dimensional Consolidation Properties of Soils (D 2435) (1996). This could underestimate the compressibility of tire chips, although the effect is less than for soils due to the flexibility of individual tire chips. Two isotropic compression tests were performed in a triaxial cell with a sample size of 71.1 mm diameter and 149 mm height. Low stress (up to 100 kPa) and high stress (up to 500 kPa) tests were conducted for both confined compression and isotropic compression. For each load increment, the load was maintained until no significant change in strain occurred. This took about 10 min for each load increment. Direct shear tests were conducted at normal stresses of 0, 20.7, 41.4, 62.1, and 82.7 kPa with a 63.5 mm diameter shear box at a displacement rate of 1 mm/min. Before the shear stress was ap-

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plied, the samples were compressed under the normal stress until there was no significant change in vertical strain. Consolidated drained triaxial tests were used to measure the strength and stress-strain response of tire chips. The samples were prepared in a rigid split-mold placed on the triaxial cell pedestal. A suction less than or equal to the planned confining pressure was applied. After removal of the mold, the triaxial cell was assembled and the confining pressure applied. The vacuum was reduced simultaneously as the cell pressure was increased. The samples were then loaded axially under a displacement-controlled mode with an axial deformation rate of 0.635 mm/min. The axial load, axial displacement, volume change, and cell pressure were recorded with a data acquisition system. Four tests were conducted with confining pressures ranging from 20 to 60 kPa. Compressibility Results Figure 1a shows that, in general, the tire chips exhibit strain hardening regardless of the testing method; however, isotropic compression produces larger volumetric strains than confined compression because of higher average stress in isotropic compression. The confined compressed specimen had about 23 and 40% volumetric strain at vertical stresses of 120 and 480 kPa, respectively. These results are similar to the confined compression results of Tatlisoz et al. (1998), where the sample strained about 26% at 120 kPa vertical stress, and Humphrey et al. (1993), where the sample strained about 40% at 400 kPa vertical stress.

Figure 1b shows that the relationship between the void ratio and the logarithm of stress for both confined compression and isotropic compression is linear above a stress of about 20 kPa. The compressibility of the tire chips may be characterized by using the slope of the void ratio versus log stress curves as a compression index: e c   log 

(2)

where e is the void ratio change and  is either the isotropic or vertical stress. The confined compression index is about 0.5 and the isotropic compression index is about 0.4. Small tire chips have much lower compressibility than largesized tire shreds (100 to 900 mm long) that exhibit 40% vertical strain at a vertical applied stress of about 17 kPa (Zimmerman 1997). One reason for this behavior is that the initial void ratio of the large-sized tire shreds was about 3, while the initial void ratio of the smaller tire chips (2 to 10 mm) was 0.98. A second reason is that individual large shreds are more compressible than individual small chips. With most other particulate systems, the solid phase is incompressible as compared with the deformations that result from particle movement. Direct Shear Test Results Figure 2 shows the shear stress and vertical deformation versus horizontal displacement for direct shear tests. None of the shear stress versus horizontal displacement curves has well-defined

FIG. 1—Compression tests of tire chips.

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FIG. 2—Direct shear tests of tire chips.

peaks up to 20 mm of horizontal displacement. This is similar to the shape of the curves presented by Humphrey et al. (1993). The shear strength of the tire chips was analyzed using MohrCoulomb theory. Because no peak shear stress was observed, shear stresses at displacements of 10, 20, and 30% of the shear box diameter were plotted versus normal stress for each test. The results are shown in Fig. 3. Generally, the stress at a displacement of 10% of the shear box diameter has been used to define the strength of the material if no peak stress occurred during shearing (AASHTO 1986). When applied to these tire chips, a friction angle ( ) of 32° with zero cohesion (c) is obtained. Using the shear stresses at 20 and 30% horizontal displacement give a of 42 and c of 5.7 kPa and a of 45° and c of 8.1 kPa, respectively. The failure envelope is curved in the low normal stress range if zero cohesion was interpreted. During shearing, the samples initially compressed and then began to dilate at about 5 to 13 mm of horizontal displacement (8 to 21% of the shear box diameter) (Fig. 2b). The volume change varied between 1.6% (compression) and 1.8% (dilation); a mini-

mum volume was reached between 8 and 21% horizontal displacement. Because the samples were forced to shear within a limited horizontal zone in the direct shear test, minimum volume represents the point where the sliding friction between particles starts to be fully mobilized. Using the minimum volume as a failure criterion, a friction angle of 41° was obtained by fitting a straight line through the data and the origin (Fig. 3). The minimum volume failure criterion provides an objective method of defining failure; however, from a design perspective, the maximum acceptable displacement may be used to define failure. The Mohr-Coulomb shear strength parameters obtained from direct shear tests by this and previous studies are shown in Table 1. Individual test results are plotted in Fig. 4 as shear stress versus normal stress. All of these results used 10% horizontal displacement (or 9% in the case of Foose et al., 1996) as the failure criterion. A power function fit to the data and through the origin (Rsquared value of 0.94) gives the following empirical equation:   1.60.75

(3)

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Friction coefficients at normal stresses up to 90 kPa can be obtained by differentiation of Eq 3: tan  1.20.25

(4)

It is concluded that the shear strength of shredded tires is independent of the particle size and that the stress range of the tests and the criterion used to define failure influence the MohrCoulomb parameters. The power function provides an empirical model for interpreting direct shear results over a wide range of normal stresses.

Direct Shear Tests on Tire Disks To better understand the characteristics of the shearing response of tire chips, the friction between two flat tire disks was measured. Ten 63.5 mm diameter disks were cut from the side walls of a used tire. Direct shear tests were conducted by shearing the two disks along their interface. The shear stress versus horizontal displacement response is shown in Fig. 5, where it is seen that the shear stress increases almost linearly up to 2.5 to 3.5 mm of displacement, and then becomes constant thereafter. A friction angle of 39° is obtained from

FIG. 3—Shear strength of tire chips from direct shear tests.

FIG. 4—Shear strength envelope of tire shreds from direct shear tests.

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tem begins to dilate, the sliding and interlocking friction is mobilized. From Fig. 2b, the tire chip specimens generally reached their minimum volume after about 15% (10 mm) horizontal displacement. The friction angle of 41° obtained at the minimum volume supports this interpretation. Triaxial Test Results

FIG. 5—Direct shear tests of tire disks.

Representative stress-strain and volume change curves of triaxial tests on the tire chips are plotted in Fig. 7. The stress-strain curves in Fig. 7a are nearly linear up to axial strains of about 15%. Beyond this strain, the curves show strain softening with clearly defined maximum deviator stresses. Figure 7b shows the volumetric strain versus axial strain curves with volume decreases up to axial strains of about 25%, where minimum volumes are reached. Beyond this strain, dilation occurs. The low-strain (up to 15%), linear portions of these curves are similar to stress-strain responses reported elsewhere (Ahmed 1993; Benda 1995; Masad et al. 1996; and Lee et al. 1999). No evidence of strain hardening was found in this study as was observed by Ahmed (1993) and Masad et al. (1996). The volume change versus axial strain for strains less than 10 to 15% is consistent with the volume change results of Benda (1995), Masad et al. (1996), and Lee et al. (1999). None of the previous studies showed the minimum volume response observed in this study. Stress-Strain Behavior Previous studies reported linear compression loading curves or slight strain hardening (Ahmed 1993; Benda 1995; Masad et al. 1996; and Lee et al. 1999). The results of this research indicate nearly linear stress strain response to 15% strains but strain softening thereafter. The linear portions of all the data on tire chips at stresses below limiting or failure stresses can be characterized with parameters analogous to Young’s modulus and Poisson’s ratio. The initial stress-strain modulus, E, is defined as: (1  3) E    1

FIG. 6—Friction angle of tire disks.

the regression of peak shear stress versus normal stress as shown in Fig. 6. The friction angle for the disks is larger than the 32° friction angle for tire chips at 10% horizontal displacement, but is smaller than 42 and 45° obtained at 20 and 30% horizontal displacement, respectively. The friction angle of 39° for the disks represents the sliding friction of rubber on rubber, and it is expected that the friction angle for tire chips would be greater due to the combined effect of interlocking of particles and sliding friction. A friction angle of tire chips that is less than the basic 39° at 10% displacement and larger than 39° at greater displacement suggests that rolling or individual particle deformation occurs in the early stage of shearing. During these small shear deformations, the system is compressing until the minimum volume is reached. At larger deformations, when the sys-

(5)

where 3 is the confining pressure, and 1 and 1 are the axial stress and strain, respectively. Linear regression on the lower strain (0 to 15%) portions of the stress-strain curves in Fig. 7a produced initial moduli of 720, 840, 870, and 920 kPa at confining pressures of 20, 28, 40, and 60 kPa, respectively. Figure 8 is a graph of the initial modulus versus confining pressure for the data from five studies in Table 2. Figure 8 illustrates that the modulus, E, increases with increasing confining pressure, but the rate of increase decreases at high confining pressures. A quadratic equation fit to the data in Fig. 8 gives: E  13.23  0.019123

(6)

A power function between confining pressure and initial modulus (Janbu 1963) was tried but rejected because the quadratic equation provided a better empirical fit. Equation 6, based on all available data, can be used to calculate the initial Young’s modulus of tire chips for confining pressures to 350 kPa and strains to 15%. Vertical strain and volume change data from the linear portions of the curves were used to calculate the lateral strain ratio, :

h   

v

(7)

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FIG. 7—Stress-strain response and volume change character of tire chips under triaxial loading.

where h is the horizontal strain, and v is the vertical strain. The strain ratios for the tests conducted here were 0.29, 0.27, 0.28, 0.28, and 0.30 for confining pressures of 20, 28, 40, and 60 kPa, respectively. There is no relationship between the strain ratios and confining pressure, so an average value of 0.28 can be used. Stress-strain parameters analogous to those in elastic theory are appropriate for the triaxial response of tire chips at strains where the curves are nearly linear; however, at higher strains where strain softening occurs, a nonlinear model is needed. The stress-strain curves from zero to maximum stress were characterized with the hyperbolic model (Duncan and Chang 1970):

1  3  

Rƒ 1   Ei (1  3)ƒ





mal strain, and, Rf is the failure ratio. Ei is the initial tangent modulus and is determined by the following equation: 3 Ei  Kpa  pa

n

 

(9)

where K is a modulus number, n is an exponent, and, pa is the atmospheric pressure expressed in the same pressure units as Ei and 3. The parameters obtained for the Duncan-Chang model are:  37°, c  0, Rf  0.453, K  10.6, and n  0.28. Shear Strength

(8)

where 1 and 3 are the major and the minor principal stress, respectively; (1 – 3)f is the stress difference at failure, is the nor-

Table 2 lists the shear strength parameters obtained from seven different triaxial test studies. All the authors conducted compression loading tests except Wu et al. (1997), who conducted compression-unloading tests. Because the stress-strain plots are linear,

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FIG. 8—Relationship between stress-strain modulus and confining pressure of tire chips in triaxial tests.

failure was defined at strains of 10 and 20%. The cohesions and friction angles from different authors vary significantly. The results from tests conducted at confining pressures less than 55 kPa have zero cohesion intercept, while tests with confining pressures up to 350 kPa exhibit cohesion up to 82 kPa. This suggests that a linear failure envelope may not be appropriate for a large range of confining pressures. The peak deviator stress observed in the tests conducted in this study is a better criterion for defining shear failure. The coincidence of peak deviator stress and minimum volume at similar axial strains suggests a failure mechanism similar to the direct shear test results. At small axial strains, individual tire chips deform and move into available void space. The strains that occur during this phase are primarily volumetric as opposed to shear strains. It is only after the minimum volume has been reached that the chips begin to shear or slide past one another. The Mohr-Coulomb parameters using maximum deviator stress as the failure criterion are a friction angle of 37° and zero cohesion. Conclusions The shear strength and stress-strain response of shredded tires has been described by a number of studies. The results of this research are synthesized with those of previous studies to draw the following conclusions with regard to the mechanical behavior of shredded tires. For direct shear tests, the stress-displacement curves were nonlinear with no well-defined peak stress for most tests. Samples compressed at low horizontal displacements reached a minimum volume, then dilated after about 15% horizontal displacement. If minimum volume is used as the failure criterion, the MohrCoulomb envelope has a friction angle of 41° with zero cohesion. Data from this and previous studies indicate that particle size does not affect the shear strength. The variation in strength parameters depends on the normal stresses at which the specimens were tested. A synthesis of all direct shear test data suggests the strength envelope for the 10% displacement failure criterion is nonlinear and described by a power function.

Both confined compression and isotropic compression tests show that the tire chips tested for this study exhibit strain hardening but the strains in isotropic compression are larger than those from confined compression. The confined compression index is about 0.5 and the isotropic compression index is about 0.4. For triaxial tests, the stress-strain response is linear up to 15% strain but exhibits strain softening beyond 15% strain. The samples decrease in volume during axial loading, and the volume change versus axial strain is almost linear up to 15% axial strain. The samples reach a minimum volume and start to dilate at strains of about 25%. The axial strain at which minimum volume occurs roughly coincides with the maximum deviator stress. If peak deviator stress, and coincidentally minimum volume, is used as failure criteria, the friction angle of the shredded rubber tested for this study is 37°. If strength and stress-strain data from all studies are considered, the unit weight and shred size do not have a significant effect on the stress-strain response or strength parameters. Based on all triaxial tests, the initial stress strain modulus increases with increasing confining pressure according to a quadratic function. For the rubber chips tested for this study, the lateral strain ratio is independent of confining pressure and averages 0.28. The stress-strain curves from zero to maximum stress were characterized with the hyperbolic model (Duncan and Chang 1970). The parameters that resulted from this exercise are:  37°, c  0, Rf  0.453, K  10.6, and n  0.28. All tests in this study were conducted on dry tire chips. In practical applications, shredded tires may be in a moist condition in the field. The effect of water on the mechanical behavior of shredded tires needs to be investigated. Acknowledgments This study was funded by grants from the Landfills Alternatives Program of the Iowa Department of Natural Resources and the University of Northern Iowa, Recycling and Reuse Technology Transfer Center. The authors would like to thank EnTire Recycling, Inc., Nebraska City, Nebraska and Four D Corp., Duncan, Oklahoma for providing the test materials. Donald T. Davidson Jr. provided invaluable assistance with triaxial testing during this study.

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References AASHTO, 1986, Standard Specifications for Transportation Materials and Methods of Sampling and Testing, Part II: Methods of Sampling and Testing, 14th ed., Washington, DC. ASTM D 6270-98, 1998, Standard Practice for Use of Scrap Tires in Civil Engineering Applications, August. ASTM D 2435-96, 1996, Standard Test Method for One-Dimensional Consolidation Properties of Soils, August. Ahmed, I., 1993, “Laboratory Study on Properties of RubberSoils,” Purdue University, Indiana, Joint Highway Research Project, Report No. FHWA/IN/JHRP-93/4. Benda, C. C., 1995, “Engineering Properties of Scrap Tires Used in Geotechnical Applications,” Report No. 95-1, Vermont Agency of Transportation, Montpelier, VT. Bosscher, P. J., Edil, T. B., and Eldin, N., 1993, “Construction and Performance of Shredded Waste Tire Test Embankment,” Transportation Research Record, No. 1345, Transportation Research Board, Washington DC, pp. 44–52. Bressette, T., 1984, “Used Tire Material As An Alternative Permeable Aggregate,” Report No. FHWA/CA/TL-84/07, Office of Transportation Laboratory, California Department of Transportation, Sacramento, CA. Duncan, J. M. and Chang, C. Y. 1990, “Non-Linear Analysis of Stress and Strain in Soils,” Journal of Soil Mechanics and Foundation Engineering, Vol. 96, SM 5, pp. 1629–1653. Foose, Gary J., Benson, C. H., and Bosscher, P. J., 1996, “Sand Reinforced with Shredded Waste Tires,” Journal of Geotechnical Engineering, Vol. 122, No. 9, pp. 760–767. Gebhardt, M. A., 1997, “Shear Strength of Shredded Tires as Applied to the Design and Construction of a Shredded Tire Stream Crossing,” MS thesis, Iowa State University. Humphrey, D. N., Sandford, T. C., Cribbs, M. M., and Manion, W. P., 1993, “Shear Strength and Compressibility of Tire Chips for

Use as Retaining Wall Backfill,” Transportation Research Record, No. 1422, National Research Council, Transportation Research Board, Washington, DC, pp. 29–35. Janbu, N., 1963, “Soil Compressibility as Determined by Oedometer and Triaxial Tests,” European Conference on Soil Mechanics & Foundation Engineering, Wiesbaden, Germany, Vol. 1, pp. 259–263. Kjartanson, B. H., Lohnes, R. A., Yang, S., Kerr, M. L., Zimmerman, P. S., and Gebhardt, M. A., 1993, “Use of Waste Tires in Civil and Environmental Construction,” Final Report, Iowa Department of Natural Resources Landfill Alternatives Financial Assistance Program. Lee, J. H., Salgado, R., Bernal, A., and Lovell, C. W., 1999, “Shredded Tires and Rubber-Sand as Lightweight Backfill,” Journal of Geotechnical and Geoenvironmental Engineering, Vol. 125, No. 2, pp. 132–141. Masad, E., Taha, R., Ho, C., Papagiannakis, T., 1997, “Engineering Properties of Tire/Soil Mixtures as a Lightweight Fill Material,” Geotechnical Testing Journal, Vol. 19, No. 3, pp. 297–304. Scrap Tire Management Council, 1999, http://www.tmn.com/ rma/html/ms5.htm. Tatlisoz, N., Edil, T. B., and Benson, C. H., 1998, “Interaction Between Reinforcing Geosynthetics and Soil-Tire Chip Mixtures,” Journal of Geotechnical and Geoenvironmental Engineering, Vol. 124, No. 11, pp. 1109–1119. Wu, W. Y., Benda, C. C., Cauley, R. F., 1997, “Triaxial Determination of Shear Strength of Tire Chips,” Journal of Geotechnical and Geoenvironmental Engineering, Vol. 123, No. 5, pp. 479–482. Zimmerman, P. S., 1997, “Compressibility, Hydraulic Conductivity, and Soil Infiltration Testing of Tire Shreds and Field Testing of A Shredded Tire Horizontal Drain,” MS thesis, Iowa State University.