Design Guidelines for Increasing the Lateral Resistance of Highway-Bridge Pile Foundations by Improving Weak Soils (2011)

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5Publications on ground improvement methods and their applications to foundation design are extensive. However, engi- neering practice typically has not used soil improvement tech- niques in combination with deep foundations to increase the lateral resistance of bridge foundations. If soil improvement has been undertaken, it has generally been with the goal of pre- venting liquefaction or increasing soil resistance so that deep foundations were not required. Three excellent publications that summarize the state of the art in soil improvement are provided by Mitchell (1981), Terashi and Juran (2000), and ASCE (1997). Based on a review of methods documented in these and other publications, specific improvement methods or technologies appropriate for applications to improve the lat- eral resistance of soils (stiffness and strength) associated with pile foundations are summarized in Tables 2-1 and 2-2. A review of the literature indicates that there are also a number of case histories that provide some insight regarding the degree of increased lateral pile resistance that could be obtained by employing soil improvement in concert with deep foundations. In connection with research studies, field load tests have been performed on two pile groups where the native clay soil was excavated and replaced with compacted granu- lar soil. These studies were primarily undertaken to evaluate group interaction factors under lateral loading. Brown et al. (1987) conducted lateral load tests on a nine-pile group in sat- urated stiff clay. Later, Brown et al. (1988) excavated the clay, compacted sand around the pile group, and repeated the lat- eral load test. Rollins et al. (2005) performed cyclic lateral load tests on a 15-pile group in medium-consistency clay. Rollins, Snyder et al. (2010) excavated the clay, replaced it with clean sand, and performed additional lateral pile group load tests. Although these tests were not designed to evaluate the effect of excavation and replacement on lateral pile group resistance, the test results can be compared to provide this information. The pile group tested by Brown et al. (1987, 1988) was a nine-pile group consisting of 0.25-m (10-in.) diameter steel pipe piles filled with grout. The piles were driven in a 3 × 3 arrangement with a 0.75-m center-to-center (3D) spacing in both directions. The original clay profile consisted of stiff, over- consolidated clay with an undrained shear strength of about 1150 psf (57 kPa) at the ground surface, which increased to about 3000 psf (150 kPa) at a depth of 18 ft (5.5 m) below the ground. Approximately 9 ft of the clay was excavated and replaced with a relatively uniform clean sand compacted to a dry unit weight (γd) of 98 pcf (15.4 kN/m3), which is a relative density of about 50%. A direct shear test indicated a friction angle of 38.5°, but back-calculated friction angles using LPILE (Reese, Wang, Isenhower et al., 2004a) suggest a friction angle of around 50°. A plot showing the total load versus deflection curves for the pile group in both clay and sand is presented in Figure 2-1. At deflections less than about 20 mm, the lateral resistance of the pile group in clay is about the same as that in sand. However, at greater deflections, the lateral resistance of the pile group in sand eventually exceeds that for the pile group in clay by over 28%, despite the fact that the clay was relatively stiff. The pile group tested by Rollins et al. (2005) and Rollins, Snyder et al. (2010) consisted of 15 12.75-in. diameter steel pipe piles driven closed-ended to a depth of about 40 ft. The piles were driven in a 3 × 5 grouping with a center-to-center spacing of 4.17 ft (3.92D) in the direction of loading and 3.5 ft (3.29D) transverse to loading. The upper 2.5 m of clay in the original soil profile had an undrained shear strength of about 900 psf. The pile group reacted against two 4-ft diameter drilled shafts. Prior to the second set of tests, the upper 1 m of clay was excavated and replaced with compacted clean sand. In addi- tion, an extra 1.5 m of sand was compacted above the original ground elevation so that the upper 2.5 of the profile consisted of clean sand compacted to 93% of the modified Proctor max- imum density. The load versus deflection curves for the pile group in clay and sand are compared in Figure 2-2. Because the clay strength was relatively soft, the lateral resist- ance of the pile group in sand was considerably higher than that for the pile group in clay. Analyses using the computer C H A P T E R 2 Available Ground Improvement Case Histories and Approaches

6Table 2-1. Summary of soil improvement methods in loose cohesionless soils. Method Advantages Disadvantages Strength Gain Cost Excavate and Replace Compacted Fill •Simple, widely used method •Equipment widely available •Compaction can be measured easily •May require excavation support or dewatering Low to mod. Low to mod. Flowable Fill •No special equipment or personnel required •No compaction necessary •Must dispose of excavated material Mod. to high Mod. to high In-Situ Densification Dynamic Compaction Densification to 25-ft depth from surface impact •Decrease in improvement with depth •Produces vibration and noise •Could produce downdrag in existing piles •Requires overhead clearance for dropping weight Low Low Vibro Compaction •Relatively uniform improvement with depth •Could produce downdrag in existing piles Low Moderate Stone Columns •Increased densification relative to vibrocompaction •Increased shear resistance of reinforced soil mass •Drainage provided by columns in the event of liquefaction concerns •Must install stone through entire layer to treat loose sand at depth •Could produce downdrag in existing piles Low Moderate Stone Columns with Wick Drains •Improved effectiveness with high fines content soils •Does not produce spoils •Increased cost and logistical effort of installing wick drains prior to stone column treatment Low Moderate Compaction Grouting •Can be used for retrofit below a pile cap •Can treat zones of interest without treating all soil above the zone •Does not produce spoils •Less effective at shallow depths where pressure is restricted •More difficult to evaluate improvement for retrofit conditions below pile Moderate Moderate Soil Mixing Deep Soil Mixing •Mixing can occur to 60- to 80-ft depths •Significant strength gain can be achieved • Can produce columns (3-ft dia.) or wall panels at desired depths •Can decrease the strength of sensitive clays •Produces spoils Mod. to high Mod. to high Grouting Permeation Grouting •Can be used for retrofit below a pile cap •Can treat zones of interest without treating all soil above the zone •Does not produce spoils •Limited to very coarse sands •More difficult to evaluate improvement for retrofit conditions below pile than in areas around periphery of pile group •Uniformity of treatment is often difficult Moderate Moderate Jet Grouting •Low noise and vibration •Can treat soil under pile cap after construction •Can transform pile group into equivalent pier for scour resistance •Creates spoil material •Requires mobilization of specialized equipment and personnel High High

7Table 2-2. Summary of soil improvement techniques in soft cohesive soils. Method Advantages Disadvantages Strength Gain Cost Excavate and Replace Compacted Fill • Simple, widely used method • Equipment widely available • Compaction can be easily measured • Must dispose of excavated material • May require excavation support or dewatering Low to mod. Low to mod. Flowable Fill • No special equipment or personnel required • No compaction necessary May need to dispose of excavated material Mod. to high Mod. to high Cement Treated Excavated Clay • Strength can be increased by 3 to 5 times the in-situ strength • Requires special care in mixing operations • Significant laboratory testing necessary to develop treatment plan • Field strength gain usually much less than laboratory because of less efficient mixing Mod. to high Mod. to high Lime Treated Compacted Clay • Strength can be increased 3 to 5 times the in-situ strength • Requires less cement to achieve same strength as using cement alone • Requires special care in mixing operations • Significant laboratory testing necessary to develop treatment plan • Field strength gain usually much less than laboratory because of less efficient mixing • Difficult to treat soils with high sulfate content Mod. to High Mod. to High Rammed Aggregate Piers • Can be constructed above or below water table • Creates extremely dense gravel columns with high-friction angle • Increases lateral pressure in surrounding soil • Increases shear resistance of the reinforced soil mass Creates relatively little increase in density of surrounding soil • Columns provide no flexural resistance. • Creates extremely dense high- friction angle columns Low Low In-Situ Soil Mixing Vibro Replacement • Increases shear resistance of the reinforced soil mass • Can cause heave of surrounding ground Low Low Compaction Grouting • Can be used for retrofit below a pile cap • Can treat zones of interest without treating all soil above the zone • Does not produce spoils • Less effective at shallow depths where pressure is restricted • Can decrease the strength of sensitive clays • More difficult to evaluate improvement for retrofit conditions below pile Moderate Low Deep Soil Mixing • Mixing can occur to 60- to 80-ft depths • Significant strength gain can be achieved • Can produce columns (3 ft dia.) or wall panels at desired depths • Can decrease the strength of sensitive clays • Produces spoils Mod. to high Mod. to high Mass Mixing • Mixing can occur in-situ to 10- to 15-ft depths • No need for excavation or recompaction • Strength increase of 3 to 5 times original shear strength • Significant laboratory testing necessary to develop treatment plan • Field strength gain usually much less than laboratory because of less efficient mixing Mod. to high Mod. to high Grouting Jet Grouting • In-situ treatment with flexibility to produce variety of geometries (columns and panels) • Flexibility to treat only zones of interest • Can treat soil under pile cap after constructio n • Can transform pile group into equivalent pier for resistance during scour events • Creates spoil material • Requires mobilization of specialized equipment and personnel High High

program GROUP (Reese, Wang, and Arrellaga et al., 2004b) were successful in matching the measured response of the pile groups in clay. However, for the pile groups in sand, successful agreement with measured response generally required the use of friction angles that are somewhat higher than would normally be used in engineering practice. The potential for improved resistance using excavation and replacement increases as the clay becomes softer and compacted granular soil becomes denser. Increases of 60% are possible (Rollins, Snyder et al., 2010). Two field test studies have evaluated the passive force on a pile cap as a function of soil type and density. Mokwa and Duncan (2001) performed tests on a 1.1-m (3.5-ft) deep and 1.9-m (6.3-ft) wide anchor block. The block was originally poured flush against an excavation into partially saturated stiff clay and a lateral load test was performed. The clay was then excavated and replaced with a compacted sandy gravel backfill and the test was repeated. Rollins, Gerber, and Kwon (2010) evaluated the passive force provided by various soils against a pile cap that was 1.1 m (3.67 ft) deep and 5.2 m (17 ft) wide. Tests were conducted on a silty sand at two densities and on loose silty sand with a 1- to 2-m wide zone of dense compacted gravel immediately adjacent to the pile cap. The native clay in the tests performed by Mokwa and Dun- can (2001) was partially saturated. Triaxial shear tests on the clay at the natural moisture content indicate that the cohesion 8 Figure 2-1. Load vs deflection curves for nine-pile groups in stiff clay and dense sand based on Brown et al. (1987, 1988). Figure 2-2. Load vs deflection curves for 15-pile group in medium stiff clay and dense sand based on Rollins et al. (2005) and Rollins, Snyder et al. (2010). 0 100 200 300 400 500 600 700 800 900 0 10 20 30 40 50 60 70 Displacement (mm) To ta l L oa d (kN ) Sand Clay 0 200 400 600 800 1000 1200 1400 1600 1800 2000 0 20 40 60 80 100 Average Group Deflection (mm) To ta l G ro u p Lo ad (kN ) Sand Clay

was 1 ksf and that the friction angle ranged from 32° to 38°. The clay was excavated to the base of the cap block and replaced by a compacted, sandy gravel. The sandy gravel (GW-GM in the Unified Soil Classification System) was com- pacted to a relative density of approximately 80%. Triaxial shear tests indicate that the friction angle could range from a low of 48° to a high of 52°. A comparison between the passive force-deflection curves for the clay and gravel is provided in Figure 2-3. In this case, the lateral resistance provided by the stiff, partially saturated clay was considerably higher than that for the gravel at the shallow depths involved. As a result, the lat- eral resistance actually decreased substantially when the com- pacted gravel was used in place of the clay. Duncan and Mokwa (2001) concluded that the log-spiral method provided the best estimate of the ultimate capacity and that the passive force- deflection relationship could be estimated reasonably using a hyperbolic curve. This conclusion was supported by analyses of additional large-scale tests by Rollins and Cole (2006) and Cole and Rollins (2006). Rollins, Gerber, and Kwon (2010) performed lateral load tests on a pile cap supported by 12 0.324-m diameter pipe piles. The piles provided sufficient vertical resistance so that the full wall friction force could develop. Basic passive force- deflection relationships were developed for two tests involv- ing silty sand compacted at 88% and 98% of the modified Proctor maximum unit weight as shown in Figure 2-4. The increased compactive effort produced a considerable increase in passive resistance. Preliminary analyses indicate that this behavior is predicted quite well using the Duncan and Mokwa (2001) approach along with the soil properties measured in the field. Tests also were performed using the loose silty sand backfill along with a well-compacted zone of sandy gravel adjacent to the pile cap. The compacted zones were 0.9 m (3 ft) and 1.83 m (6 ft) thick. These tests indicate that compacting rela- tively thin zones (3 to 6 ft wide) of sandy gravel around a pile cap can significantly increase the passive resistance as illus- trated in Figure 2-5. In this case, replacing a 0.9 m (3 ft) zone of loose silty sand around the pile cap with compacted gravel increased the lateral resistance on the pile cap from an initial value of 70 kips to more than 180 kips, which is an increase of over 200%. Crack patterns from the tests, shown in Figure 2-6 indicate that the compacted gravel zones increase the effective 9 Figure 2-3. Comparison of passive force provided by stiff, partially saturated clay and compacted sandy gravel against 1.1-m deep  1.9-m wide cap block. Figure 2-4. Measured passive force vs deflection relationships for two full-scale tests with silty sand compacted to 88% and 98% of the modified Proctor maximum unit weight. Clay Gravel 0 100 200 300 400 500 600 700 0 10 20 30 40 50 Displacement (mm) To ta l L oa d (kN ) 0 50 100 150 200 250 300 350 400 0 0.5 1 1.5 2 2.5 3 Deflection (inches) Pa ss iv e Fo rc e (ki ps ) Dense Silty Sand (98%) Loose Silty Sand (88%)

width of the pile cap and reduce the pressure on the loose silty sand behind it, thereby increasing passive resistance. Rollins and Nasr (2010) completed a follow-up study to develop a gen- eralized equation to predict the increased passive resistance that could be obtained by constructing a narrow dense granu- lar zone adjacent to a pile cap or abutment surrounded by loose sand. This investigation followed the same pattern employed in this NCHRP study. The full-scale load tests reported by Rollins, Gerber et al. (2010) and Gerber et al. (2010) were used to calibrate soil parameters within the finite element computer program PLAXIS. In addition, the computer model was veri- fied against analytical solutions for computing passive force (e.g., Duncan and Mokwa, 2001). Once the FEM analysis model was calibrated, parametric studies were performed to evaluate the influence of changes in wall height (H), the dense zone width (Bf), the friction angle of the dense zone (φD) and the friction angle of the looser surrounding sand (φL). Fig- ure 2-7 shows the layout of the improved zone relative to the pile cap or abutment and defines the basic parameters involved. To facilitate generalization, the width of the dense zone is nor- malized by the height of the pile cap. In addition, the ultimate passive resistance provided by the limited-width dense granu- lar backfill (PLW) is normalized by the ultimate passive resist- ance for a homogeneous dense backfill (PFW) that fully encloses the passive failure surface. This ratio is defined as the passive force ratio (PFR). PFR is plotted as a function of normalized width (Bf/H) for a number of cases in Figure 2-8 where the dense zone consisted of gravel. As the normalized width increases, the PFR increases. When Bf/H exceeds 1.0, the PFR is normally greater than 60%, and the PFR increases as the fric- tion angle of the surrounding sand increases. Typically, the Bf/H ratio would need to be around 4.0 to fully enclose the fail- ure surface; therefore, the narrow dense zone is relatively effec- tive in mobilizing the majority of the total passive resistance. Based on the result from the FEM parametric study, Equation 1 10 0 50 100 150 200 250 300 0 1 2 Deflection (inches) Lo ad (k ip s) Loose Silty Sand/6 ft Gravel Backfill Loose Silty Sand/3 ft Gravel Backfill Loose Silty Sand Figure 2-5. Passive force vs deflection curves for loose silty sand against a 17-ft wide by 3.67-ft high pilecap after excavation and replacement with 3-ft and 6-ft zones of compacted gravel backfill against the cap. (a) 3 ft sandy gravel zone plus loose silty sand backfill (b) 6 ft sandy gravel zone plus loose silty sand backfill 10 ft x 17 ft x 3.67 ft Pile Cap 2 ft x 2 ft Grid 10 ft x 17 ft x 3.67 ft Pile Cap 2 ft x 2 ft Grid Gravel Backfill Loose Silty Sand Figure 2-6. Plan view of crack patterns behind a pile cap after excavation and replacement of loose silty sand with (a) a 3-ft and (b) a 6-ft zone of compacted sandy gravel behind the pile cap. Figure 2-7. Layout of geometry for limited width dense backfill zone adjacent to pile cap or abutment. Loose Sand (φL) Pile Cap or abutment D en se S an d Direction of Lateral Load H (φ D ) Bf 2 ft min.

was developed to predict the PFR for a limited-width dense granular backfill. At a test site on Treasure Island in San Francisco Bay, stone column treatment was used to improve a 6-m-thick liquefi- able sand layer around two test foundations (Ashford, Rollins, and Baez 2000a; Ashford, Rollins, Bradford et al., 2000b; and Weaver et al., 2005). One test foundation was a 0.6-m dia- meter cast-in-steel shell (CISS) shaft that reacted against a group of four 0.324-m diameter driven steel pipe piles. Lateral load tests were first performed on the test foundations prior to treatment, then comparable lateral load tests were performed after treatment for comparison. A high-speed hydraulic actu- ator was used to apply load. In addition to the conventional static lateral load tests, a pattern of small explosive charges was detonated sequentially to produce a liquefied volume of soil within which the test foundations also could be laterally loaded to large displacement levels (9 in.). The layout of the test foundations is shown in Figure 2-9. The 0.9-m diameter stone columns were installed in a square pattern with a spac- ing of 2.5 m from center to center (see Figure 2-9) using the dry, bottom feed method. The sand had a mean grain size of about 0.2 mm with a fines content between 5% and 10% and was initially placed using hydraulic filling techniques. Prior PFR B HD D f= − − + ( ) − 3 1418 0 139 0 033 0 484 0 043 . . . . . φ φ B H B Hf L L f( ) + − ( )2 20 003 0 007 1. . ( )φ φ 11 Figure 2-8. Percentage of total passive resistance developed by a narrow dense gravel zone adjacent to an abutment relative to that for a homogenous dense gravel zone. 0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100% 0 0.5 1 1.5 2.52 PF R = P L W /P F W G ra v e l BF/H ϕ(gravel)=39° ϕ(sand)=36° ϕ(gravel)=39° ϕ(sand)=32° ϕ(gravel)=39° ϕ(sand)=27.7° to treatment, correlations with SPT and CPT test results (Kulhawy and Mayne, 1990) indicated that the sand had a relative density of about 50%; however, after treatment, the relative density was increased to between 85% and 100%. A comparison of the load versus deflection curves for the CISS pile and the pile group before and after treatment but prior to blast liquefaction are shown in Figure 2-10. The lateral resistance was increased by about 25% for the pile group and about 33% for the CISS pile. Back-analyses using the computer program LPILE indicate that the friction angle was increased from about 39° prior to treatment to about 48° after treatment. Prior to stone column treatment the sand layer liquefied (excess pore pressure ratios of 100%) following blasting and remained liquefied for at least 10 minutes. Settlement fol- lowing dissipation of excess pore pressures amounted to about 12 in. A plot of the load-deflection curve for the CISS pile following blast-induced liquefaction is provided in Fig- ure 2-11(a). The pre-blast load-deflection curve also is shown for comparison. Following liquefaction, about 11 times more movement was required to develop the same lateral resist- ance as that prior to liquefaction. After treatment, high excess pore pressure developed initially after the blasting but dissi- pated within a few seconds, presumably due to the increased drainage provided by the stone columns. The load-deflection curve after blasting is presented in Figure 2-11(b) for compar- ison and the stiffness of the curve is several hundred percent higher in comparison to that prior to treatment.

12 x xx x x x x x x x xx x x xx o oo o o o o o o o o o o o x = Charges for first and second blast @ 3.5 m o = Charges for second blast only @ 3.5 m = Stone Column, 0.9 m Diameter, Installed to a depth of 6 m = 2 x 2 Pile Group = 0.6 m CISS Pile Limit of Excavation Limit of Densification Scale: 1m Figure 2-9. Layout of test foundations, stone columns, and explosive charges.

13 -100 0 100 200 300 400 -20 -10 0 10 20 30 40 50 Displacement (mm) Lo ad (k N) Post-Treatment Pre-Treatment (a) Pile Group (b) CISS Pile -100 0 100 200 300 400 -20 -10 0 10 20 30 40 50 Displacement (mm) Lo ad (k N) Post-Treatment Pre-Treatment Figure 2-10. Static load vs deflection curves before and after stone column treatment for (a) pile group and (b) 0.6-m CISS pile. Figure 2-11. Pre- and post-blast cyclic load-deflection curves (a) before and (b) after stone column treatment for a 0.6-m CISS pile. Pre-Blast Post-Blast (b) After Stone Column Treatment -200 -100 0 100 200 300 400 500 600 700 -25 0 25 50 75 100 125 150 175 200 225 250 Displacement (mm) Lo ad (k N) -200 -100 0 100 200 300 400 500 600 700 -25 0 25 50 75 100 125 150 175 200 225 250 Displacement (mm) Lo ad (k N) Non-Liquefied Liquefied (a) Before Stone Column Treatment