Current design codes of Mechanically Stabilized Earth (MSE) Walls allow the use lower lateral earth pressure coefficient (K value) for designing geosynthetics walls than those used to design steel walls. The reason of this is because geosynthetics walls are less rigid permitting the wall to deform enough to work under active pressures instead of at rest pressures as in steel walls. A new concept of crimped steel bars was recently introduced. This new type of bar was tested for tension and pullout behavior. Results on tests made on crimped bars show that putting those crimps in the steel bar will give us a better pullout behavior and a more flexible tensile behavior. This new type of steel bar will behave more like geosynthetics, allowing the wall to deform sufficiently to reach the necessary deflection to reach the active condition. The use of steel by current design codes is pushing MSE walls to be designed with more steel than needed. Measurements of the force in different walls showed that the steel is not being used even close to the maximum stress allowed by the code which is 50%. The proposed design methodology using crimped bars will help us save around 52% of steel volume compared to the actual design procedures. This means a huge improvement in the usage of steel versus actual designs. This improvement is obtained because of the efficient behavior of rounded bars under corrosion and because of the flexibility in the bars obtained with the crimps that will allow us to reach the active condition.

Large scale static lateral load tests were completed on a pile cap with wingwalls under several different sand backfill configurations: no backfill, loosely compacted unconfined, loosely compacted slip plane wall confined, loosely compacted MSE wingwall confined, and densely compacted MSE wingwall confined. The relative compaction of the backfill was varied during each test to observe the change in passive resistance provided by the backfill. The wall types were varied to observe the force placed on the walls and the wall displacement as a result of the laterally loaded pile cap and backfill relative compaction. Passive force-displacement curves were generated from each test. It was found that the densely compacted material provided a much greater passive resistance than the loosely compacted material by 43% (251 kips) when confined by MSE walls. The outward displacement of the MSE walls decreased noticeably for the dense MSE test relative to the loose MSE test. Backfill cracking and heave severity also increased as the relative compaction level of the backfill increased. As the maximum passive force was reached, the reinforcement reached their peak pullout resistance. Correlations were developed between the passive pressure acting on the pile cap and the pressure measured on the MSE wingwalls as a function of distance from the pile cap for both loose and dense backfills. The pressure measured on the wingwalls was approximately 3 to 9% of the pressure acting on the pile cap. As the distance from the pile cap increased, the pressure ratio decreased. This result helps predict the capacity of the wingwalls in abutment design and the amount of allowable wall deflection before pullout of the backfill reinforcement occurs. Three methods were used to model the measured passive force-displacement curves of each test. Overall, the computed curves were in good agreement with the measured curves. However, the triaxial soil friction angle needed to be increased to the plane strain friction angle to accurately model both the loose and dense sand MSE and slip plane wall confined tests. The plane strain friction angle was found to be between 9 to 17% greater than the triaxial friction angle.

Large scale static lateral load tests were performed on a pile cap under varying sand backfill configurations: no backfill, full-width dense sand backfill, dense sand slip plane confined backfill, and two configurations of dense sand MSE wall confined backfills. Efforts were made to maintain the relative compaction of the backfills for each of the tests near the same value. The MSE wall panel arrangement was varied to determine the effect of different reinforcement configurations on the passive resistance and wall panel displacement. Passive force-displacement curves were generated from each test. It was found that the MSE design manual provided reasonable estimates of pullout resistance of bar mats in dense sand, and that the passive resistance of a soil backfill confined by MSE walls can be calculated with an increased friction angle using a log-spiral approach. Also, the amount the triaxial friction angle can be increased depends on how much the MSE wall panels displace outward. Correlations were developed between the pressure on the pile cap and that on the MSE wall panels near the pile cap. Generally, the pressure on the wall panels was less than 10% of that which was on the adjacent pile cap, and decreased as the distance from the pile cap increased. Finally, it was found that while limiting the backfill width decreases the ultimate passive resistance of the backfill, if the backfill is confined in a plane strain configuration the passive resistance per unit width is higher than that for an unconfined backfill.

Full scale lateral load tests were performed on five piles located at various distances behind MSE walls. Three of the five test piles were production piles used to support bridges, and the other two piles were located behind a MSE wing walls adjacent to the bridge abutment. The objective of the testing was to determine the effect of spacing from the wall on the lateral resistance of the piles and on the force resisted by the MSE reinforcement. Tentative curves have been developed showing p-multiplier vs. normalized spacing behind wall for a length to height ratio of 1.1 and 1.6. The data suggest that with a L/H ratio of 1.6, a p-multiplier of 1 can be used when the normalized distance from the back face of the MSE wall to the center of the pile is at least 3.8 pile diameters. When the L/H ratio decreases to 1.1 a p-multiplier of 1 can be used when the pile is at least 5.2 pile diameters behind the wall. A plot showing the induced load in the reinforcement as a function of distance from the pile has been developed. The data in the plot is normalized to the maximum lateral load and to the spacing from the wall to the pile. The best fit curve is capped at a normalized induced force of approximately 0.15. The data show that the induced force on the reinforcement when a lateral load is applied to the piles decreases exponentially as the normalized distance from the pile increases. The plot is limited to the conditions tested, i.e. for the reinforcement in the upper 6 ft. of the wall with L/H values ranging from 1.1 to 1.6.

Full scale lateral load tests were performed on four piles located at various distances behind MSE walls. Three of the four test piles were production piles used to support bridges, and the other pile a production pile used as part of the bridge abutment. The objective of the testing was to determine the effect of spacing from the wall on the lateral resistance of the piles and on the force resisted by the MSE reinforcement. Lateral load-displacement curves were developed for pile at various spacing and with various reinforcement ratio (reinforcement length, L divided by wall height, H). The force in the reinforcement was measured using strain gauges. Lateral load analyses were performed to determine the minimum spacing required to eliminate any effect of the wall on the pile resistance (p-multiplier of 1) and the reduction in soil resistance at closer spacings (p-multiplier less than 1). With the addition of the data fro Price (2012) tentative curves have been developed showing p-multiplier vs. normalized spacing behind wall for a length to height ratio of 1.6, 1.2, and 1.1. The data suggest that with a L/H ratio of 1.6, a p-multiplier of 1 can be used when the normalized distance from the back face of the MSE wall to the center of the pile is at least 3.8 pile diameters. When the L/H ratio decreases to 1.2 and 1.1 a p-multiplier of 1 can be used when the pile is at least 4.5 and 5.2 pile diameters behind the wall respectively. For smaller spacings, the p-multipliers decreased essentially linearly with normalized distance from the wall. A plot showing the increased load in the reinforcement as a function of distance from the pile has been developed. The data in the plot is normalized to the maximum lateral load and to the spacing from the wall to the pile. The best fit curve is capped at a normalized tensile force of approximately 0.12. The data show that the increase in tensile force on the reinforcement when a lateral load is applied to the piles decreases exponentially as the normalized distance from the pile increases. The plot is limited to the conditions tested, i.e. for the reinforcement in the upper 3 ft. of the wall with L/H values at 1.2.

A full scale MSE wall was constructed and piles were driven at various distances behind the wall. Lateral load testing was conducted and the performance of the pile, wall, and reinforcement were measured. The piles were 12.75 inch pipe piles, and the wall was reinforced with welded wire grid reinforcement. The objective of the testing was to characterize the relationship between the lateral pile resistance and the distance of the pile behind the back face of the MSE wall. Load-displacement curves are presented for the piles located behind the wall at 66 inches (5.3 diameters), 55 inches (4.3 diameters), 41 inches (3.2 diameters), and 24 inches (1.9 diameters). The lateral resistance of the piles decreases as the spacing behind the wall decreases. The results of the testing have been matched in LPILE using p-multipliers to reduce the lateral resistance. A curve has been developed showing the variation of p-multiplier with normalized pile spacing behind the wall, including data from previous studies. The curve suggests that a p-multiplier of 1 (no reduction in lateral resistance) can be used when the normalized distance from the back face of the wall to the center of the pile is at least 4 pile diameters. The p-multiplier decreases relatively linearly for smaller spacings.

ABSTRACTLateral Resistance of Piles near 15 Foot Vertical MSE AbutmentWalls Reinforced with Ribbed Steel StripsJarell Jen Chou HanDepartment of Civil and Environmental Engineering, BYUMaster of ScienceA full scale MSE wall was constructed and piles were driven at various distances behind the wall. Lateral load tests were conducted to determine the effect of pile spacing from the wall on the lateral resistance of the piles and the force resisted by the MSE reinforcement. The piles used for this study were 12.75 inch pipe piles and the reinforcements were ribbed steel strips.Load-deflection curves were developed for piles located behind the wall at 22.4 inches (1.7 pile diameters), 35.4 inches (2.8 pile diameters), 39.4 inches (3.1 pile diameters) and 49.9 inches (3.9 pile diameters). Data results show that the lateral resistance of the pile decreases as the spacing behind the wall decreases. Measured load-deflection curves were used to compare with computed curves from LPILE with p-multiplier developed for the lateral resistance of piles closer to the wall. A curve was created showing the variation of p-multiplier with normalized pile spacing behind the wall. The curve suggests that a p-multiplier of 1 (no reduction in lateral resistance) can be used when a pile is placed at least four pile diameters from the back face of the wall.

The basic mixture of lightweight cellular concrete (LCC) consists of cement, water, and a stable foaming agent. It is generally classified as having a density of less than 50 pounds per cubic foot (pcf), which is less than both traditional concrete and backfill materials. LCC has gained popularity in construction due to its lightweight, self-leveling, and ease of production and placement. These characteristics have made LCC a popular lightweight backfill material for mechanically stabilized earth (MSE) walls. However, there has been relatively little research on the large-scale behavior of LCC as a MSE backfill. Therefore, large-scale test results defining failure mechanisms and the strength criteria of reinforced LCC are extremely valuable. In this study, a three walled test box (10 ft wide x 12 ft long x 10 ft high) was constructed to contain the LCC. Two 5 ft tall x 10 ft wide MSE wall segments were poured and cured, before being placed as the fourth wall of the test box. The test box was built with a steel reaction frame to reduce lateral deflections during testing of the LCC that was not in the direction of the MSE wall, thus creating a two-dimensional or pseudo "plane strain" geometry. The box was filled with four lifts of Class II LCC 2.5 feet thick with ribbed-strip reinforcements at the center of each lift. After the LCC was cured, two static load tests were performed by applying surcharge to the surface of the LCC using six hydraulic jacks. The static load tests compared the LCC behavior of an MSE wall in comparison with unreinforced LCC without MSE wall panels. Multiple forms of instrumentation were used to understand the behavior of the LCC during surcharge loading. The instrumentation also helped to characterize the strength criteria for LCC based on failure in the large-scale and laboratory testing. This was done to determine the failure mechanism for the MSE wall retaining system with ribbed-strip reinforced LCC backfill. Data was gathered primarily through lateral wall pressures, lateral wall deflections, and forces induced on the ribbed-strip reinforcements. The test results show that an MSE wall with LCC backfill can withstand significant surcharge loading with limited axial and lateral deformations. However, failure occurred at surcharge pressures of about 60% of the unconfined compressive strength. The use of a retaining system significantly increased the failure loads and produced a more ductile failure mode than Class II LCC with a free-face. The active pressures observed are similar to a granular material with a friction angle (ϕ) of 34°, Ka=0.28, and a cohesion of 700 to 1600 psf for Class II LCC. Likewise, failure of the free-face occurred at a value predicted by Rankine theory with ϕ = 34° and c = 1600 psf.

Researchers performed 35 full-scale lateral load tests on piles driven within the reinforcement zone of a mechanically stabilized earth wall (MSE wall). Data defining the induced tensile force on the reinforcements during lateral pile loading was used to develop multi-linear regression equations to predict the induced tensile force. Equations were developed by previous researchers that did not consider the diameter of the pile, the fixed head condition, relative compaction, or cyclic loading. The purpose of this research was to include this tensile force data and develop prediction equations that considered these variables. Additionally, a full-scale lateral load test was performed for a 24-inch diameter pipe pile loaded against a 20-inch width square pile. The test piles were instrumented using load cells, string potentiometers, LVDTs, strain gauges and hybrid pressure sensors. The lateral load tests were used to evaluate the ability of finite difference (LPILE) and finite element (PLAXIS3D) models to compute results comparable to the measured results. The results of this analysis showed that the diameter of the pile is a statistically significant variable for the prediction of induced tensile force, and the induced tensile force is lower for piles with larger diameter. Fixed head conditions have no effect on the prediction of induced tensile force. Cyclic loading had minimal impact on the prediction of induced tensile force, but relative compaction did have an important statistical significance. Therefore, prediction equations for induced tensile force in welded wire were developed for relative compaction less than 95 percent and relative compaction greater or equal than 95 percent. A general prediction equation (Eq. 3-4) was developed for ribbed-strip reinforcements that included the effect of pile diameter and larger head loads. With 1058 data points, this equation has an R2 value of 0.72. A general prediction equation (Eq. 3-9) was also developed for welded-wire reinforcements that included data from cyclic and static loading, fixed and free head conditions, and relative compaction for 12-inch wide piles with a higher range of pile head loads. This equation based on 2070 data points has an R2 value of 0.72. The prediction equations developed based on all the available data are superior to equations developed based on the original set of field tests. The finite element models produced results with good agreement with pipe pile measurements while the finite difference model showed better agreement with the square pile measurements. However, for the denser backfills involved, back-calculated soil properties were much higher than would be predicted based on API correlations. The API equations are not well-calibrated for dense granular backfills.

The purpose of this study was to extract galvanized steel wire reinforcement coupons from mechanically stabilized earth (MSE) walls along I-15 and determine the rate of corrosion that has taken place since Phase I, which was conducted by Gerber and Billings (2010). The galvanized steel reinforcement analyzed in this study has been in place for 19 to 20 years at the time of extraction. A total of 85 coupons were extracted and laboratory analysis was performed to determine the thickness of remaining zinc galvanization on each coupon. Soil samples were obtained from each one-stage wall extraction location to determine moisture content for correlation with corrosion. After laboratory testing was performed, the measured zinc coating thickness was compared to that determined in Phase I. An average corrosion rate of approximately 0.032 oz/ft²/year has occurred since Phase I. According to the AASHTO (2017) design corrosion rate of 0.35 oz/ft²/year for the first two years and 0.09 oz/ft²/year until the depletion of the zinc, the zinc coating would have been completely depleted after 16 years. Based on the results of laboratory testing, the initial galvanization coating was likely greater than the specified thickness of 2.0 oz/ft² (86 μm). The zinc galvanization is corroding at a slower rate than the AASHTO design rate. The AASHTO design rate for depletion of zinc coating and subsequent corrosion of the steel reinforcement is conservative for the corrosion conditions present in the MSE wall reinforcement coupons tested. The integrity of the steel reinforcement that is currently in place is not likely to be compromised by corrosion.