3.2.1 “Mini Pier” Experiments (Adams et al., 2002, 2007)

Five “mini pier” experiments were conducted and reported by Adams and his colleagues at the Turner‐Fairbank Highway Research Center, Federal Highway Administration (Adams et al., 2002, 2007). The test specimens were cuboidal in shape (referred to as “mini piers”), measuring 1.0 m by 1.0 m in cross‐section and 2.0 m in height. The mini piers were subjected to incrementally increasing vertical loads on the top surface under an unconfined condition. Figure 3.3(a) shows a mini pier experiment after it has been loaded to failure. Of the five experiments, one was conducted without reinforcement, two were on reinforcement spacing of 0.2 m, and the other two on mixed reinforcement spacing of 0.4 m to 0.6 m. Two woven geotextiles of different stiffness/strength were used as reinforcement, with wide‐width strengths (T_f ) of 70 kN/m and 21 kN/m.

The stress–strain curves of the five mini piers are shown in Figure 3.3(b). It is seen that the reinforced piers, especially Piers C and D on reinforcement spacing of 0.2 m, exhibit much higher stiffness and strength than the unreinforced pier (Pier A). The relative influence of reinforcement spacing vs. reinforcement strength on stress–strain behavior can be evaluated by inspecting carefully the difference in stress–strain relationships between Pier D (on 0.2 m spacing) and Pier B (on 0.4 m spacing), as compared to the difference between Pier D (T_f = 70 kN/m) and Pier C (T_f = 21 kN/m). It is obvious that the former is much greater than the latter, indicating that the influence of reinforcement spacing is greater than that of reinforcement strength. For a quantitative evaluation, the applied vertical pressures needed to bring about a global vertical strain of 1.0% are 190 kPa, 245 kPa, and 300 kPa for Piers B, C, and D, respectively. Considering Pier D as a baseline for comparisons, the required vertical pressure to reach 1.0% strain for Piers B and C is, respectively, 1.6 times and 1.2 times that for Pier D, while the reinforcement strength of Pier C is 3.5 times that of Pier D, yet the spacing of Pier B is only twice that of Pier D. This series of experiments provide solid evidence that reinforcement spacing plays a more important role than reinforcement strength in the load–deformation behavior of reinforced soil.

3.2.2 Unconfined Compression Experiments (Elton and Patawaran, 2005)

Elton and Patawaran (2005) conducted seven field‐scale unconfined compression experiments on 0.76 m diameter, 1.5 m high cylindrical specimens (see Figure 3.4). Reinforcement spacing was 0.15 m for all the experiments, except one with 0.3 m. The reinforcements were needle‐punched nonwoven geotextiles of different weights, designated TG500, TG600, TG700, TG800, TG1000, and TG028, with respective ultimate strength of 9, 14, 15, 19, 20, and 25 kN/m, as shown in Table 3.1. The soil was a poorly graded sandy soil with less than 5% fines and a maximum grain diameter of 12.7 mm. Modified Proctor compaction tests (ASTM D1557) of the soil gave an optimum moisture content of 9.3% and a maximum dry unit weight of 19.0 kN/m³. Direct shear tests on the sand in a moist compacted state (at a wet unit weight of 18.5 kN/m³ ) gave an angle of internal friction (ϕ) of 40° with cohesion (c) of 29 kPa.

The stress–strain curves of the seven soil–geosynthetic composites are shown in Figure 3.5. It is seen that reinforcement strength influences the stress–strain behavior appreciably. The curves follow a general trend, namely, the stronger the reinforcement, the higher the strength of the soil–geosynthetic composite. The experiment designated TG500‐12, the only one on reinforcement spacing of 0.3 m, shows much lower composite strength than TG500 which has the same reinforcement as TG500‐12, but on half the reinforcement spacing. The influence of reinforcement spacing is seen to be very significant.

3.2.3 Generic Soil–Geosynthetic Composite Plane‐Strain Experiments (Wu et al., 2010, 2013)

Wu et al. (2010, 2013) conducted and reported the results of four generic plane‐strain soil–geosynthetic composite (SGC) experiments and, for comparison purposes, one soil‐only experiment (see Figure 3.6). Since a soil mass with layers of geosynthetic inclusion is far from being uniform, it was deemed necessary that the test specimen be sufficiently large to provide “representative” behavior of soil–reinforcement composites. The finite element method of analysis with an elasto‐plastic hardening soil model was used to determine the specimen dimensions needed to provide representative stress–strain–volume change behavior of the soil–geosynthetic composites. Specimens of different dimensions were analyzed, and the global (overall) stress–strain and volume change relationships are shown in Figure 3.7. It was concluded that a specimen 2.0 m in height and 1.4 m in depth on 0.2 m spacing under a confining pressure of approximately 30 kPa would give an adequate representation of 7.0‐m high GRS on 0.2 m reinforcement spacing. The conditions of the five experiments are summarized in Table 3.2. Among the five SGC experiments, all except one were conducted under a confining pressure of 34 kPa, and the experiments varied in reinforcement stiffness/strength and reinforcement spacing.

The backfill used in the experiments was a well‐graded gravelly sand. The Mohr–Coulomb strength parameters for as‐compacted soil are: c = 70 kPa and ϕ = 50° for confining pressure between 0 and 200 kPa; c = 240 kPa and ϕ = 38° for confining pressure between 200 and 750 kPa. The reinforcement was a woven geotextile with a wide‐width strength of 70 kN/m, except for Test 3, where the reinforcement was formed by gluing two sheets of the same geotextile together, producing twice the stiffness and the wide‐width strength of a single sheet (as verified by wide‐width tensile tests) while maintaining the same soil–reinforcement interface conditions as for a single sheet.

The measured load–deformation and volume change relationships of the soil–geosynthetic composites obtained from the SGC experiments are shown in Figure 3.8. A summary of the test conditions and ultimate vertical load‐carrying pressure of the SGC experiments is given in Table 3.3. The following observations are made:

• Using vertical strains at failure as an indicator for the ductility of the soil–geosynthetic composites, all soil–geosynthetic composites (Test 2 to Test 5) exhibit much higher ductility than the unreinforced soil mass (Test 1). The vertical strain at failure of Test 2 (the baseline) is more than 200% of Test 1.
• All the soil–geosynthetic composites (Test 2 to Test 5) have significantly higher load‐carrying capacity than the unreinforced soil (Test 1). The ultimate load‐carrying capacity of Test 2 is about 350% of Test 1.
• Test 2 (s_v = 0.2 m) and Test 3 (s_v = 0.4 m) have the same T_f /s_v ratio, yet the load‐carrying capacity of Test 3 is only 65% of Test 2.
• The influence of reinforcement spacing can be seen by comparing the load–deformation curves of Test 4 and Test 2. By reducing reinforcement spacing by 50% (and maintaining the same reinforcement strength), the load‐carrying capacity is seen to increase about 100%. The ductility is also seen to reduce significantly with the increase in reinforcement spacing.
• The influence of reinforcement strength can be seen by comparing the load–deformation behavior of Test 3 and Test 4. By reducing the reinforcement strength by 50% (and maintaining same reinforcement spacing), the load‐carrying capacity is seen to decrease only about 25%. Comparing this to the influence of reinforcement spacing (see the bullet immediately above), it is evident that reinforcement spacing has a much stronger influence than reinforcement strength on the load‐carrying capacity and load–deformation behavior of GRS.
• The load‐carrying capacity of Test 5 (unconfined) is about 30% lower than that of Test 2. This shows that confining pressure can have a significant influence on the load–deformation behavior of GRS.

3.3.1 Mechanisms of Apparent Confining Pressure and Apparent Cohesion

If we have a bucketful of dry clean sand and wish to make it into the steepest possible pile without external support or compaction, we can perhaps accomplish it by pouring the sand onto a flat surface very slowly. We will then likely find the steepest angle to be between 30° and 40°. This angle is commonly referred to as the angle of repose. Without cohesion between dry granular particles, the steepest angle that any granular fill can assume is less than about 50°. Only with cohesion can a granular fill stand at a 90° angle without lateral support.

Figure 3.9(a) shows a clean dry sand at its angle of repose, and Figure 3.9(b) shows the same sand reinforced by strips of plain paper embedded horizontally inside the sand. By incorporating layers of reinforcement in the sand, the sand is internally reinforced so that it can assume a 90° angle. The embedded paper strips are said to generate an apparent cohesion in the dry sand and allow the sand to behave as if it possesses a certain amount of cohesion. Note that the paper is wrapped at the face to prevent loss of dry sand particles (sloughing); it otherwise assumes little structural function at the face.

The mechanism of apparent cohesion as a reinforcing mechanism for reinforced soil was first identified by Schlosser and Long (1974). In this mechanism, the interaction between soil and tensile reinforcement is said to bring about an apparent cohesion in the soil without introducing any appreciable time‐dependent deformation commonly associated with cohesive soils. As seen in the Mohr circles shown in Figure 3.10(a), under a confining pressure σ_3c , the apparent cohesion acquired by tensile reinforcement, c_R , will lead to an increase in shear strength, i.e., an increased major principal stress at failure from σ₁ to σ_1R. In the figure, it is tacitly assumed that the internal friction angle of the soil ϕ will not be affected by the presence of tensile reinforcement. Limited data available to date suggest that the internal friction angle of geosynthetic reinforced soil will indeed stay approximately the same as that of the unreinforced soil (see Figure 3.11). Some recently published data seems to contradict this assumption. Additional reliable test data from field‐scale experiments (such as those described in Section 3.2) are needed before a conclusion regarding the assumption can be reached.

Yang (1972) proposed a reinforcing mechanism, known as the mechanism of apparent confining pressure, which is somewhat parallel to the mechanism of apparent cohesion. In the mechanism of apparent confining pressure, the reinforcement is said to generate an added confining pressure in the direction of the reinforcement layers. The added confining pressure can be regarded as an increase in the minor principal stress, Δσ_3R , and will lead to a larger major principal stress at failure, σ_1R , as seen in Figure 3.10(b).

The mechanisms of apparent cohesion and apparent confining pressure are interrelated. By equating the shear strength of the two mechanisms (i.e., setting σ_1R of a given σ_3c to be equal for the two mechanisms), see Figure 3.10, the apparent cohesion of a reinforced soil can be determined from the apparent confining pressure, and vice versa. This is addressed below.

For a reinforced soil mass of which the reinforcement has a tensile rupture resistance of T_f , and vertical spacing between adjacent layers of reinforcement of s_v , it has been assumed that the increase in confining pressure due to tensile inclusion, Δσ_3R , is:

(3‐1)

By equating the σ_1R in Figure 3.10(a) to the σ_1R in Figure 3.10(b), and making use of Rankine’s passive pressure theory, the apparent cohesion, c_R , can be expressed as:

(3‐2)

Substituting Eqn. (3‐1) into Eqn. (3‐2), the apparent cohesion becomes

(3‐3)

where K_P is the coefficient of Rankine passive earth pressure. Note that Eqn. 3‐3 is for a cohesionless soil, c = 0.

Eqn. (3‐3), if valid, can be very useful for the stability analysis of earth structures involving reinforced soil. Given the strength parameters of soil in the unreinforced zone (c = 0 and ϕ > 0) and T_f  and s_v of the reinforcement in the reinforced zone, Eqn. (3‐3) can be used to determine the strength parameter c_R of the reinforced soil as a function of T_f and s_v . By assuming ϕ_R = ϕ (i.e., the same friction angle for unreinforced and reinforced soils), the stability analysis of a reinforced soil structure can be carried out the same way as for soil‐only earth structures by using c = 0 and ϕ for soil in the unreinforced zone, and c = c_R and ϕ for soil in the reinforced zone.

Moreover, with the aid of Eqn. (3‐3), the ultimate load‐carrying capacity, q_ult , of a reinforced granular soil mass (c = 0) becomes:

(3‐4)

In recent years, however, the validity of Eqns. (3‐1) to (3‐4) has been brought into question. The seemingly correct expression of Δσ_3R = T_f /s_v (i.e., Eqn. (3‐1)) suggests that the reinforcing effect is reflected completely by the simple ratio of T_f and s_v . It implies that the resulting reinforcing effect is the same for any two reinforcing schemes of the same T_f /s_v ratio. In other words, reinforced soil walls of any value of reinforcement spacing will behave the same as long as the reinforcement strength values are proportional to the reinforcement spacing. This assumption has encouraged designers to use larger reinforcement spacing with proportionally higher strength reinforcement because the cost of higher strength geosynthetic products does not differ very significantly, yet placing reinforcement on larger spacing would save considerable time in field installation. Consequently, reinforcement spacing as large as 0.6–1.0 m (or even larger) has routinely been employed in geosynthetic reinforced soil walls since manufacturers have little difficulty producing ever stronger reinforcement. It has been common practice for GMSE walls (MSE walls with geosynthetics as reinforcement) to use reinforcement on 0.3–0.9 m or 12–36 in. spacing (with 0.6 m or 24 in. being typical) to reduce construction time.

Field‐scale experiments, such as the FHWA mini pier experiments (see Section 3.2.1) and the SGC experiments (see Section 3.2.3), have demonstrated conclusively that reinforcement spacing plays a more significant role in the load–deformation behavior of a reinforced soil mass than reinforcement strength. Wu et al. (2010) and Wu and Pham (2013) have presented an analytical model which introduces a W‐factor in the expression for Δσ_c to reflect more accurately the relative roles of reinforcement spacing (s_v) and reinforcement strength (T_f ) as:

(3‐5)

where d_max = maximum particle size of the soil. Eqn. (3‐5), referred to as the W‐equation, leads to a revised expression for apparent cohesion (c_R):

(3‐6)

where c = cohesion of unreinforced soil. It follows that the ultimate load‐carrying capacity of a soil–geosynthetic composite, q_ult , becomes:

(3‐7)

Eqn. (3‐7) is referred to as the load‐carrying capacity equation of soil–geosynthetic composites.

For GRS walls with dry‐stacked concrete block facing, the value of σ_c can be estimated as , where γ_b is the bulk unit weight of the facing block, D is the depth of the facing block unit (in the direction perpendicular to the wall face), and δ is the friction angle between vertically adjacent facing blocks (or between the geosynthetic reinforcement and the facing block, if the geosynthetic is sandwiched between adjacent blocks). The value of σ_c is very small for light‐weight dry‐stacked concrete blocks without any connection enhancement elements (e.g., lips, keys, or pins), and can be conservatively assumed to be zero.

The load‐carrying capacity equation, Eqn. (3‐7), has been shown to give drastically improved prediction of load‐carrying capacity of soil–geosynthetic composites than Eqn. (3‐4) (Wu et al., 2010, 2013). The Federal Highway Administration (FHWA) compared all available measured data where a GRS mass is loaded to failure with those calculated by the load‐carrying capacity equation (Eqn. 3‐7), as seen in Figure 3.12. The GRS included in the comparison are field‐scale experiments with soils ranging from fine sand to crushed gravel, geosynthetic reinforcement varying from light‐weight nonwoven to heavy‐weight woven geotextiles, and reinforcement spacing ranging from 0.15 m to 0.3 m. It is seen that the load‐carrying capacity equation, Eqn. (3‐7), provides rather accurate predictions of the load‐carrying capacity of GRS for all the cases. The FHWA (Adams et al., 2011a,b) has incorporated Eqn. (3‐7) in the design guidance of a GRS bridge system called the geosynthetic reinforced soil–integrated bridge system (GRS‐IBS).

3.3.2 Mechanism of Suppression of Soil Dilation

When subject to shear stress, a dense granular soil tends to dilate if volume change is permitted. Leonards (1962) used a set of simple drawings (Figure 3.13(a)) to explain the dilative behavior. In a dense granular soil, soil particles packed tightly and will need to roll over past other particles (without breaking) to produce shear‐induced deformation and lead to an increase in volume, i.e., dilation. If dilation is suppressed, i.e., if shear deformation is prohibited, the soil stiffness and strength will increase.

Suppression of soil dilation in reinforced soil may be a result of two effects. The first is illustrated conceptually in Figure 3.13(b) where the reinforcement in tension contains soil particles to stay between reinforcement layers and prevent them from rolling over past neighboring soil particles, hence suppress the increase in volume. The second effect is related to restraint of lateral deformation due to soil–reinforcement interface friction (see Section 3.3.3). The latter effect is explained below.

When a soil mass is subjected to vertical loads, there will be a reduction in volume in the vertical direction (i.e., contraction) and an increase in volume in the lateral direction (i.e., expansion). In a dense granular soil, the latter typically overcomes the former, and results in a net increase in volume, hence dilation. For reinforced soil, however, the increase in volume in the lateral direction is significantly smaller than that in unreinforced soil, which can result in smaller dilation. For reinforced soil with closely spaced reinforcement, the dilative lateral volume change can be less than the compressive vertical volume change. This can lead to complete suppression of dilation or even a net decrease in the total volume.

Suppression of soil dilation by the presence of reinforcement layers can be seen by inspecting Figure 3.14, which shows measured stress–stain and volume change relationships of unreinforced and reinforced soils obtained from the field‐scale SGC plane‐strain experiments described in Section 3.2.3. In the unreinforced soil, dilation (i.e., increasing volumetric strain due to shear), after a certain amount of initial contraction, is seen from measured data. In the reinforced soils, on the other hand, the dilative behavior is clearly seen to be suppressed due to the presence of geosynthetic reinforcement.

Depending on soil type, reinforcement spacing, reinforcement stiffness, and, to a lesser degree, confining pressure, the presence of geosynthetic reinforcement may or may not be able to fully suppress soil dilation. The volume change behavior due to change in reinforcement spacing is shown in Figure 3.15. The influence of reinforcement spacing is obvious. The unreinforced soil is seen to exhibit significant dilation. For reinforced soil with s_v = 1.0 m, soil dilation still occurs, but is much less than that of unreinforced soil; with s_v = 0.5 m, the soil–geosynthetic composite becomes nearly incompressible (i.e., no net volume change) once the axial strain exceeds about 2%; with s_v = 0.2 m, dilation of the composites is suppressed over the entire range of stress and strain; the tendency for suppression of dilation is even more pronounced for s_v = 0.1 m than for s_v = 0.2 m.

The dilation angles (denoted by ψ) of the soil–geosynthetic composites for reinforcement spacing of 0.1 m, 0.2 m, 0.5 m, 1.0 m, and unreinforced soil are about –10°, –3.5°, 0°, +3°, and +20°, respectively (a smaller negative ψ‐value indicates more contraction, implying less tendency to dilate). It is conceivable that ψ‐value may be used to characterize the extent of suppression due to geosynthetic reinforcement. As reinforcement spacing decreases from ∞ (unreinforced) to 0.1 m, the corresponding dilation angle ψ gradually decreases from +20° to –10°, with s_v = 0.5 m showing nearly incompressible behavior.

It is interesting to note that tests of small‐size specimens have been seen to produce entirely different behavior. Figure 3.16 shows triaxial test results of a sand with zero, one, two, and three layers of geosynthetic reinforcement. The test specimens used in the tests were quite small (100 mm in diameter and 200 mm in height), and reinforcement spacing ranged from 50 mm to 100 mm. The presence of reinforcement is seen to increase the stiffness and strength of the soil, but does not offer any suppression effect on the dilative behavior observed in the unreinforced soil. In fact, the tests show that the dilative behavior is more prevalent with decreasing reinforcement spacing. Similar findings of triaxial tests performed on small size specimens have been reported, including the tests conducted by Broms (1977), which have been cited frequently in the literature. When the material in a test specimen is non‐uniform, as in a soil–geosynthetic composite, it is important to bear in mind that the test specimen needs to be large enough to accurately account for the interaction among the different components and to give representative composite global behavior, otherwise it may lead to misleading results, as exemplified by the behavior seen in Figure 3.16.

3.3.3 Other Reinforcing Mechanisms

Increase in Compaction‐Induced Stress

When a laterally confined soil mass is subjected to external vertical loading, there will generally be an increase in both vertical and lateral stresses in the soil mass. If the loading is subsequently removed, the increase in vertical stress will essentially vanish, but not the increase in lateral stress. The net increase in lateral stress is commonly known as the residual lateral stress or lock‐in lateral stress. For the same vertical loading, the stronger the restraint to lateral deformation, the higher the residual lateral stress. The residual lateral stress resulting from fill compaction operations, which typically involves a series of loadings and unloadings onto the top surface of a compaction lift, is referred to as compaction‐induced stress (CIS).

Compaction‐induced stress is an important factor to consider in terms of the load–deformation behavior of a compacted fill. Compaction‐induced stress in unreinforced soil has been the subject of many studies, including Rowe (1954), Broms (1971), Aggour and Brown (1974), Seed (1983), Bolton (1990), and Duncan et al. (1991). These studies have indicated that compaction‐induced stress can be rather significant if there is sufficient restraint to lateral movement of the soil, and will lead to increased stiffness and strength.

When a soil mass is reinforced by layers of geosynthetic reinforcement, compaction‐induced stress is more pronounced than in an unreinforced condition because a reinforced soil has a higher degree of restraint to lateral deformation of the soil due to soil–reinforcement interface friction. For the same reason, closer reinforcement spacing tends to result in higher compaction‐induced stress. Closer reinforcement spacing also “compels” the compactor operator to keep compaction lift thickness small (no more than reinforcement spacing) and improves the efficiency of fill compaction.

Figure 3.17 shows a simplified stress path at a given depth in the fill due to multiple passes of compaction. The stress path assumes that the slopes of unloading and reloading are the same. The slopes of the stress path can be calibrated by data obtained from well‐controlled field‐scale experiments of GRS, and the value of CIS can then be evaluated. Wu and Pham (2010) have provided an example.

Restrain Lateral Deformation of Soil

When a soil mass is compressed in one direction (usually the vertical direction), tensile strains tend to develop in the orthogonal direction (usually the horizontal direction). Reinforcement placed in the expected direction of tensile strains of a GRS mass will help restrain the tensile deformation of the soil mass, hence bringing about greater compaction‐induced stress, and improve stiffness and strength of the soil. The restraint of tensile strain is derived from interface bonding between the soil and reinforcement. This restraining behavior can be seen by comparing the deformation of two triaxial test specimens when subject to vertical loads: soil only vs. soil with one layer of reinforcement at the mid‐height (see Figure 3.18). The restraining effect of lateral deformation by reinforcement is evident in the photos.

Knowing the distribution of lateral strains in an unreinforced earth structure can be helpful in determining locations where reinforcement will produce the greatest improvement. This can be accomplished by performing finite element analysis on the unreinforced earth structure to identify the locations of tensile strains. The most beneficial orientation of geosynthetic reinforcement may not be the horizontal direction. However, placing geosynthetic sheets horizontally is usually least time‐consuming in actual field installation.

Stabilization of Potential Failure Wedge

If and when a potential failure wedge is to develop within the reinforced zone of a GRS wall, the geosynthetic reinforcement extending beyond the failure wedge is said to act as frictional quasi‐tiebacks to help stabilize the potential failure wedge, as seen in Figure 3.19. This failure mechanism has been considered as a major mode of failure for checking the internal stability (i.e., failure occurred within reinforced zone) of GMSE walls, and has been used for the determination of the required reinforcement length of GMSE walls.

Preserve Soil Integrity

Many nonwoven and woven geotextiles can serve the function of filtration well, and allow water to pass through without compromising the integrity of the soil. This mechanism can help maintain the stability of GRS walls in certain situations. As an example, wrapped‐face geotextile reinforcement (see Figure 3.20) has been found to help alleviate bridge bumps (differential settlement found at the junction between a bridge deck and the fill behind) by the Wyoming Department of Highways (Edgar et al., 1989). A major reason for the ability to eliminate bridge bumps by geosynthetic inclusion was not the geotextile’s reinforcing function, rather it was attributed to the capability of the wrapped‐face geotextile to contain soil particles in the approach fill. The wrapped‐face geotextile was able to prevent soil particles from falling into the gap formed between the abutment wall and the fill due to shrinkage of the bridge deck in cold seasons (Wu and Helwany, 1990; Monley and Wu, 1993).

Acceleration of Porewater Pressure Dissipation

Some geosynthetics, notably heavier needle‐punched geotextiles, can serve the function of transmitting water in the plane of their matrices in addition to reinforcing. These geosynthetics, when used in clayey fill, can facilitate drainage and accelerate dissipation of excess porewater pressure, hence allowing the soil to gain strength with time more rapidly. In Japan, for example, clayey soil is occasionally used as backfill for construction of embankments and walls due to the wide availability of some clayey soils such as a volcanic ash soil known as Kanto loam; the mechanism of accelerating porewater pressure dissipation is important in this application.

Figure 3.21 shows a clay embankment reinforced by a nonwoven needled geotextile with different layout on the two sides. It is seen that the side with more layers of geotextile settles visibly less (i.e., stronger) than the other side with fewer geotextile layers. Measurement of porewater pressure has verified the capability of the geotextile to accelerate dissipation of the porewater pressure (Yamauchi, 1987). Since the stiffness of most nonwoven needled geotextiles with good drainage capability is usually too low to serve effectively as reinforcement, geotextile manufacturers in Japan have produced woven–nonwoven composite geotextiles to combine the advantages of the two types of geotextile, i.e., good capability to dissipate the porewater pressure of a nonwoven needled geotextile, coupled with the higher tensile stiffness of a woven geotextile.

Improve Ductility of Soil Mass

Figure 3.22 shows the stress–deformation relationships of a ductile material versus a brittle material with about the same strength. Ductility is a term used to denote a material’s ability to undergo significant deformation before rupture. The term “ductility” also applies to a structure. For design of any structure constructed in public places, it is highly desirable that the structure possesses sufficiently high ductility so that if the structure is to fail unexpectedly (due to unexpected external influence or deficiencies in design or construction), failure will not be sudden catastrophic collapse.

Unreinforced soil structures with granular soils often exhibit low ductility. The lack of ductility, however, can be improved significantly with the inclusion of closely spaced geosynthetic reinforcement. This is seen in full‐scale laboratory experiments and in actual earth structures. The increase in ductility can be evaluated by examining the stress–deformation curves of field‐scale experiments of reinforced versus unreinforced soil. In the SGC plane‐strain experiments, for example, the strain at failure of GRS is about 200% of that in an unreinforced condition (see Figure 3.8). In the unconfined FHWA mini pier experiments, the ratio of failure strains between GRS and unreinforced soil is much higher, between 4.5 and 12.5 (see Figure 3.3(b)).

3.4.1 Wrapped‐Face GRS Wall

A geotextile reinforced soil wall with a wrapped face was first constructed in Siskiyou National Forest in Oregon in 1974, and soon after in Olympic National Forest in Shelton, Washington in 1975 by the U.S. Forest Service (Steward and Mohney, 1982). The satisfactory performance and low cost of these walls provided an impetus for many wrapped‐face GRS walls subsequently constructed in the U.S. and around the world. Most wrapped‐face GRS walls have used geotextiles as reinforcement and as facing. More recently, use of geogrid as reinforcement with a pre‐seeded mat or geocomposite as facing has seen applications.

Wrapped‐face GRS walls constructed to date are of two types: (a) geotextile wrapped‐face GRS walls and (b) wire‐mesh wrapped‐face GRS walls. Figure 3.24 shows three completed wrapped‐face GRS walls; the first is a geotextile wrapped‐face wall, and the other two are wire‐mesh wrapped‐face walls. The differences between the two wall types are described in Section 6.1.2. The face of wrapped‐face walls is sometimes covered with shotcrete, gunite, or asphalt emulsion (see Figure 3.25) to prevent deterioration (due to extended UV exposure) and vandalism. The construction procedure for wrapped‐face GRS walls is given in Section 6.1.2.

The backfill used in the construction of wrapped‐face GRS walls is typically granular soil, ranging from silty sand to coarse gravel. Compacted cohesive backfill has also been used, especially in Asia and South America. Both nonwoven needle‐punched and heat‐bonded geotextiles and woven geotextiles have been employed as reinforcement. Wrapped‐face GRS reinforced soil walls have been constructed mostly in remote areas or as temporary walls; however, some have been for permanent applications in an urban setting. They have also been used commonly as the “stage‐1” wall in construction of a “two‐stage wall”, where the 2^nd stage involves installing a permanent‐looking wall façade. Wrapped‐face GRS walls generally range in height from 1 m to 6 m, with taller walls being built mostly in Asia, where their faces are often vegetated. Figure 3.26 shows a very tall (21‐m high) vegetated face GRS wall in Japan and a partially vegetated face in Canada.

Experiences distilled from construction of wrapped‐face GRS walls include the following:

• Wrapped‐face GRS walls are very economical and can be built by a general contractor.
• Wrapped‐face GRS walls do not exhibit any appreciable creep as long as the backfill is predominantly granular and is well‐compacted.
• The geosynthetic reinforcement in a wrapped‐face GRS wall can effectively generate apparent cohesion (see Section 3.3.1) for a cohesionless fill to assume a vertical slope while subject to common traffic loads. Figure 3.27 shows a photo of an unprotected vertical cut in a wrapped face wall five years after the cut was made.
• Wrapped‐face GRS walls can tolerate large differential settlement without visible distress.
• The facing of wrapped‐face GRS walls may experience significant deformation, especially if a weak reinforcement is employed or if the fill is not compacted well.
• A wrapped GRS wall face covered with asphalt products or shotcrete is not aesthetically appealing compared to other types of wall face.

Most wrapped‐face GRS walls constructed to date have been designed based on the U.S. Forest Service method (see Section 5.4). Observation and field measurements of wall performance reveals that the design method is very conservative.

3.4.2 Concrete Block GRS Wall

Since being introduced in the mid‐1980s, concrete block reinforced soil walls have quickly become the most popular type of reinforced soil walls in North America. The configuration of a concrete block geosynthetic reinforced wall is seen in Figure 3.23. Figure 3.28 shows four completed concrete block GRS walls, including (a) a single tier concrete block wall transitioned into a three‐tier concrete block wall in DeBeque Canyon, Colorado, (b) a concrete block wall with its highest point standing 16.7 m (55 ft) tall in Grand County, Colorado, (c) a concrete block wall of multiple block colors in a rural setting, Winter Park, Colorado, and (d) a concrete block wall with a curved face in an urban setting, Centennial, Colorado.

The facing of concrete block GRS walls is typically formed by dry‐stacking concrete blocks, which are usually made light and small enough to be handled by a worker with bare hands. The blocks may be dry‐ or wet‐cast machine molded, with the former being far more common. A variety of split face concrete masonry units (CMUs) and segmental retaining wall (SRW) units varying in size, shape, weight, color and texture are available commercially (see Figure 3.29). The most common dimensions of split face CMUs used in the construction of GRS walls are 410 × 200 × 200 mm (16 × 8 × 8 in.). SRW units, on the other hand, are concrete blocks specifically made for segmental retaining walls, typically ranging from 100 to 200 mm (4 to 8 in.) in height and 200 to 750 mm (8 to 30 in.) in width. CMU blocks cost about $1.50 per block, depending on the region and the cement content. SRW blocks, on the other hand, typically cost between $4.50 and $6.50 per block.

Concrete blocks are either stacked vertically or at a small batter (usually less than 10° degrees from the vertical) to form the face of a concrete block GRS wall. Most concrete blocks have built‐in hollow cores/cells which may be filled with crushed stone or sand/gravel during construction to increase block weights. Some blocks have cast‐in lips or keys to help with facing alignment and to increase the bond strength between blocks. Geosynthetic reinforcement is usually sandwiched between vertically adjacent blocks through interface friction to connect the reinforced soil mass with the wall face. Some wall builders also install mechanical shear pins between blocks to help maintain facing alignment. The facing connection stability of concrete block walls is discussed in detail in Section 5.3.5.

When concrete block walls were first introduced to the wall industry, they were used mostly for landscaping, with typical wall height less than 2 m (6 ft). The use of geosynthetic inclusion, however, took the wall system to new heights because geosynthetic inclusion (especially when deployed on tight spacing) strengthens the backfill and reduces the lateral earth pressure. Concrete blocks, while providing a certain degree of local bending resistance, can be merely an architectural façade in an internally stabilized wall system (see opening statement of this chapter). Concrete block GRS walls typically range from 2 to 9 m (6 to 30 ft) in height, although walls over 20 m (65 ft) in height have been constructed without problems. Concrete block GRS walls are commonly built with a slightly battered face, although some wall builders consider the batter of little practical value. When the total height of a wall exceeds about 6 m (20 ft), most concrete block GRS walls are built in tiers.

In addition to the advantages for all GRS walls noted in Section 3.1, concrete block GRS walls enjoy the distinct advantage of being rather aesthetically pleasing. A combination of a concave face and the rough textured finish (“split face”) of concrete blocks (see Figure 3.28(a) to (c)) can help mask minor construction misalignment. Both CMUs and SRW units of different colors, textures, shapes, and sizes can be used to improve the aesthetic appeal of a GRS wall (see Figure 3.28). CMU blocks can be cut to fit or the block tails broken off to accommodate tight curves. Specially‐made narrow SRW units also serve the same purpose.

Material and engineering costs for concrete block GRS walls typically range from $5 to $10 per square foot of wall face (or $50 to $100 per square meter of wall face), excluding the cost of fill material. Not unlike other GRS walls, judicious use of on‐site soils may be permitted for construction of concrete block GRS walls, although it is generally not recommended especially if the on‐site soil is cohesive or poorly drained. Construction of concrete block GRS walls is addressed in Section 6.1.1.

3.4.3 Cast‐in‐Place Full‐Height Facing GRS Wall

Beginning in the mid‐1980s, Fumio Tatsuoka at the University of Tokyo and researchers at the Japan Railway Technical Research Institute jointly developed a GRS wall system with cast‐in‐place full‐height rigid facing (Tatsuoka et al., 1992, 1997). This system is herein referred to as the cast‐in‐place full‐height facing (CIP‐FHF) GRS wall system.

Figure 3.30 shows the detailed construction sequence of a CIP‐FHF system. The construction involves two major stages. In the 1^st stage, a wrapped‐face GRS wall, with layers of geosynthetic inclusion approximately 0.3H to 0.4H (H = wall height) in length on 0.3 m spacing (typical), is erected using stacked gravel‐filled gabions as temporary facing. This temporary wrapped‐face GRS wall is allowed to deform under its own weight (Step 5, Figure 3.30). In the 2^nd stage, full‐height rigid concrete facing (cast‐in‐place) is constructed over the temporary face of the GRS wall (Step 6, Figure 3.30). The rigid concrete facing is securely attached to the reinforced soil mass via protruded steel anchors that are embedded in the reinforced soil mass during the 1^st stage construction (Step 4, Figure 3.30). Figure 3.31 shows photos of CIP‐FHF systems soon after the 1^st stage construction, in‐service condition, and 15 years after being in service.

When rigid facing is employed in a GRS wall, a major problem is with mobilization of loads in the reinforcement, as the rigid facing will restrain lateral deformation and inhibit the reinforcing function. This problem is resolved by the CIP‐FHF system, which involves a two‐stage construction procedure. Reinforcement loads in the system are mobilized during 1^st stage construction before installation of cast‐in‐place facing in the 2^nd stage.

Since the temporary GRS wall in the 1^st stage construction has deformed under its self‐weight, the lateral earth pressure exerted on a completed cast‐in‐place full‐height facing is negligible. This allows the wall system to withstand much greater subsequent external loads, including seismic loads due to earthquakes. The rigid facing, in combination with the tight connection between the reinforced soil mass and the facing, also serves to increase lateral stress in the soil mass hence increase stiffness and strength of the soil behind the facing.

Due to its superior capability to withstand earthquakes and prolonged heavy rainfalls, the CIP‐FHF system became the default retaining wall type in Japan Railway’s wall repertoire. Figure 3.32 shows a CIP‐FHF GRS wall before and after the 1995 Kobe earthquake, which measured 7.9 on the Richter scale. The wall’s capability to survive major earthquakes is evident. As of 2008, over 120 km of CIP‐FHF walls had been built for railway embankments in Japan with great success (Tatsuoka, 2008). In other parts of the world, however, the CIP‐FHF system has seen few applications, primarily because of cost considerations and longer construction time.

3.4.4 Precast Full‐Height Panel Facing GRS Wall

In 1996, the Colorado Department of Transportation (CDOT) developed a new reinforced soil wall system with full‐height precast panel facing, referred to as an independent full‐height facing (IFF) reinforced soil wall. An IFF reinforced soil wall system, as shown schematically in Figure 3.33, comprises three major components: (i) precast full‐height reinforced concrete facing panel, (ii) reinforced soil mass, and (iii) flexible face anchor. The face anchor serves to connect the first two components, i.e., facing panels and reinforced soil mass, into an integral system. In this system, the connection needs to allow outward movement of the facing panel to take place during construction (hence, referred to as “flexible” face anchors). In the absence of this feature, the reinforcement in the reinforced soil mass would hardly be able to mobilize any loads, and would serve hardly any reinforcing function. Also, since outward movement of the facing panel is allowed during construction, a mechanism for adjusting the final positions of the panels must be in place so that the panels can be aligned in the final phase of the construction.

An IFF reinforced soil wall, locally known as the Fox wall, was constructed in 1996 along the entrance ramp from Interstate‐25 to Interstate‐70. The wall was an extension of an existing cantilever reinforced concrete retaining wall, of which the facing had the appearance of full‐height grooved concrete. For aesthetic reasons, the GRS wall must bear the same appearance as the existing wall, hence full‐height facing was needed. The total cost of an IFF reinforced soil wall system was estimated to be about half the cost of a conventional reinforced concrete cantilever wall on piles. The decision to go with the IFF reinforced soil wall system, however, was not based on cost, rather it was because little or no excavation would be needed for construction of an IFF wall system. This is important for two reasons. First, the wall was situated along a very busy section of Interstate‐25, so minimal excavation meant minimal traffic disruption. Second, the subsoil on the project site was found to have been contaminated in the past. Excavation of the contaminated soil, if needed, would have involved laborious laboratory testing and securing permits, and would likely have resulted in undue delay. Since construction, the Fox wall has performed very satisfactorily.

To construct an IFF reinforced soil wall, the facing panels are first erected at a slight batter in a shallow trench with the aid of temporary bracing. The reinforced soil mass is then constructed behind the facing with inclusion of tensile reinforcement layers on prescribed spacing. Face anchors are installed at pre‐selected elevations to attach facing panels to the reinforced soil mass. The temporary bracing is removed after the facing panels have been securely attached to the reinforced soil mass. The final step of construction involves adjusting the positions of the facing panels to achieve proper alignment. Construction of IFF reinforced soil walls is discussed in more detail in Section 6.1.3. Figure 3.34 shows two photos of the Fox wall during construction. Figure 3.35 shows the Fox wall in a service condition.

The IFF reinforced soil wall system is under‐used in the repertoire of reinforced soil walls. The IFF system offers a few distinct advantages over conventional cantilever reinforced concrete retaining walls, including the following:

• The full‐height facing gives the appearance of a strong and permanent structure.
• Construction of IFF reinforced soil walls is relatively rapid; it does not require any on‐site concrete formwork or piles.
• Construction of IFF reinforced soil walls requires no over‐excavation, hence little traffic disruption during construction.
• IFF reinforced soil walls are low in cost. The total cost of an IFF GRS wall is comparable to that of a concrete block GRS wall, and is about half the cost of a conventional cantilever reinforced concrete retaining wall on piles.

3.4.5 Timber Facing GRS Wall

A GRS wall system with timber facing was developed in the early 1980s by the U.S. Forest Service. In areas where timber logs can be obtained economically, timber facing GRS walls may be the least expensive wall type. Figure 3.36 shows two completed timber facing GRS walls, one in an urban setting and the other in a rural setting. The facing of the GRS wall is made of wooden timber logs that serve as temporary forming and also as final facing. All the timber logs need to be preservative‐treated, and are inter‐connected behind the facing by forming elements (e.g., plywood boards) with nails, as shown in Figure 3.37. Geosynthetic reinforcement is attached to the facing between forming elements and timber logs, and with a “tail” folding flat toward the back of wall. The inter‐connected timber facing offers both local and (small) global bending stiffness. Construction of timber facing GRS walls is discussed in Section 6.1.4.

Timber facing GRS walls that have been constructed to date are typically less than 6 m (20 ft) in height. Higher walls are commonly built in tiers, which allow room for planting in the setback area and enhance aesthetics. A wide variety of soils, ranging from granular soil to low‐quality clayey soil, have been used as backfill. Various geosynthetics, including woven geotextiles, nonwoven geotextiles, and geogrids, have been used as reinforcement.

Timber facing GRS walls have been researched rather extensively. In 1991, two 2.8‐m (9.2‐ft) high timber facing geotextile reinforced soil walls (known as the Denver test walls), one with a cohesive backfill and the other a granular backfill, were constructed and load‐tested under well‐controlled conditions in the laboratory. The reinforcement was a light‐weight (3 oz/yd²) nonwoven heat‐bonded polypropylene geotextile on 0.28‐m (11‐in.) vertical spacing (Wu, 1992a,b). The test walls were featured in the International Symposium on Geosynthetic‐Reinforced Soil Retaining Walls, in which researchers were invited to predict wall behavior. The material properties, construction, measured behavior, and predictions of the two walls were documented (Wu, 1992c). A few highlights of the test wall are as follows:

• With a very weak geotextile as reinforcement, the surcharge pressure at failure for the granular‐backfill wall was 200 kPa (29 psi), and for the cohesive‐backfill wall was slightly over 230 kPa (33 psi). Available design methods for GRS walls predicted surcharge pressures at failure (i.e., for factor of safety = 1.0) ranging from 0 (i.e., the wall cannot be erected) to 50 kPa (7.3 psi) for the granular‐backfill wall, much lower than the measured value (Claybourn and Wu, 1992). None of the design methods allows cohesive backfill. Additional details about how measured behavior compared to design methods as well as their implications are given in Section 5.2.
• The cohesive‐backfill wall is at least as strong as the granular‐backfill wall provided that the moisture of the soil is kept relatively constant (at 2% wet of optimum in this case).
• Neither wall exhibited appreciable creep deformation under 103 kPa (15 psi) surcharge, typical traffic load used in design.

Timber facing geosynthetic reinforced walls have shown satisfactory performance in actual construction. Wu (1994) reported applications of timber facing GRS walls in Colorado, including the use of double‐sided walls as rock fall barriers developed by wall builder Bob Barrett.

Timber facing GRS walls offer a number of advantages over other types of GRS walls, including the following:

• With timber being a natural material, timber facing GRS walls have the distinct advantage of a natural appearance, as compared to GRS walls with concrete facing.
• Where there is abundant supply of timber logs, a timber facing GRS wall is usually the least expensive wall type.
• Construction of timber facing GRS walls is simple and rapid; there is no requirement for an external forming system for construction. The wall can be built by a general contractor.
• Timber facing GRS walls can withstand fairly large settlement and differential settlement without hindering their function.
• The lifespan of the timber facing tends to dictate the service life of the wall system. When the timber is properly treated, the design life of a timber facing wall can be around 50 years.

3.4.6 Other Types of GRS Walls

In addition to the five wall types described in Sections 3.4.1 to 3.4.5, a number of reinforced soil walls (both GMSE walls and GRS walls) with different facing systems have been constructed. Examples of such walls are rock facing GRS walls, articulated panel facing GRS walls, tire facing GRS walls, welded wire mesh facing GRS walls, and gabion facing GRS walls, as shown in Figures 3.38 to 3.42, respectively. Note that the latter two wall types are sometimes constructed as the first‐stage wall of a two‐stage wall system, with the second stage often being concrete panel facing.

3.5.1 Wrapped‐Face GRS Walls

Advantages

Wrapped‐face GRS walls have demonstrated the following advantages over conventional reinforced concrete walls:

• Construction of wrapped‐face GRS walls is rapid and requires no heavy construction equipment.
• Wrapped‐face GRS walls are inherently flexible, hence can tolerate larger differential settlement without visible distress.
• When properly designed and constructed and with well‐compacted granular fill, wrapped‐face GRS walls have high load‐carrying capacity (see Figure 3.43, where a wrapped‐face GRS composite with bed sheets as reinforcement and a road base as fill material is seen to be able to carry 22 “Jersey barriers” made of solid concrete; no time‐dependent deformation was observed), and exhibit high ductility (i.e., can deform considerably without sudden collapse).
• Wrapped‐face walls do not require embedment into the foundation soil, thus eliminate the need for sub‐excavation. This feature is especially important when environmental constraints are present. Potential damage to adjacent steams and root systems due to excavation can also be avoided to protect natural environments.
• High‐quality granular backfill, although highly preferred, is not mandatory for the construction of wrapped‐face GRS walls.
• A wide variety of geosynthetics with different stiffness/strength and costs are readily available and can usually be easily transported to remote construction sites; geosynthetics are free from corrosion problems and generally have strong resistance to bacterial action.
• Other than reinforcing, geotextiles can serve multiple functions, e.g., facilitate drainage of fill material, preserve soil integrity, and maintain separation of different soils during construction and under repeated external loading (see Section 4.1 for further discussion).
• The overall cost of wrapped‐face GRS walls is low. In many applications, it has proven to be the lowest cost of all reinforced soil walls.

Disadvantages

Wrapped‐face GRS walls have a number of disadvantages compared to conventional reinforced concrete retaining walls:

• Wrapped‐face GRS walls lack the appearance of a permanent structure. When the face is covered with asphalt products or shotcrete, it is less than aesthetically pleasing.
• Wrapped‐face GRS walls may experience appreciable deformation, especially if a weak geotextile reinforcement is employed or if the fill is not well compacted.
• Some geotextiles are susceptible to chemical degradation and may deteriorate when exposed to UV lights for an extended period of time (say, more than a few days).

3.5.2 Concrete Block GRS Walls

Advantages

In addition to all the advantages of wrapped‐face geotextile reinforced walls described in Section 3.5.1, concrete block GRS walls offer a few advantages:

• A concrete block wall face is durable and can be aesthetically appealing. Concrete block GRS walls generally give the perception of a more “permanent” structure.
• A wide variety of concrete blocks with different sizes, shapes, weights, textures, and colors are readily available as facing units of GRS walls (see Figure 3.29).
• Concrete block GRS walls can be easily adapted to situations where fairly sharp curves in the longitudinal direction are needed or desired.
• Since concrete blocks offer high local compressive and some bending resistance, the deformation of concrete block GRS walls is typically much smaller than that of wrapped‐face GRS walls.
• When very small reinforcement spacing (say 0.1 m or 4 in.) is used, it promotes better compaction near the face of the wall hence improved performance near the wall face, which is especially important for GRS bridge abutments (Adams, personal communication).

Disadvantages

Concrete block GRS walls can overcome most of the disadvantages of wrapped‐face GRS walls described in Section 3.5.1, especially problems with aesthetics and deformation. However, close attention and control of facing block placement is required for concrete block GRS walls to avoid misalignment or uneven batter of the wall face. Effloresence (mineral salts leached from soil and concrete to the surface; see Figure 3.44) will leave stains on a concrete wall face if not properly treated. In addition, concrete block facing may experience problems with deterioration within a few years after it is put into service. Figure 3.45 shows an example of severe block deterioration. The cause of such deterioration, which has been seen to occur in both SRW blocks and CMU blocks, may be multi‐fold, including insufficient cement content, poor quality control, reactive aggregates, etc. Adherence to specifications for the use of concrete blocks as facing units of GRS walls is needed. Moreover, proper control of surface runoff to direct the runoff away from the wall face is a key requisite.

3.5.3 Timber Facing GRS Walls