Design and Load Testing of Large Diameter Open-Ended Driven Piles (2015)

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38 chapter five CASE HISTORIES During the course of the literature review and the interviews with agency and private-sector entities, several case histories of uses of LDOEPs were obtained. Many of the case histories included load testing of LDOEPs. Table 2 lists the case his- tories included in this chapter with basic information for each. This chapter illustrates successful testing and use of LDOEPs, as well as some lessons learned by owners, design- ers, and contractors. As such, the case histories included here are meant to be a select sample of the many case histories and reports available in the published literature. A complete list of all available case histories was not within the scope of this report. It is important that the reader be aware that the summaries in this report are brief by design and cannot include many of the details of how tests were interpreted. The results presented here are those presented by the authors of the papers reporting each project. The reference for each case history can be consulted for details on test interpretation, site conditions, etc. As mentioned earlier in this report, one issue among state DOTs is selecting resistance factors for design. Most of these case histories did not include much or any discussion on resistance factor selection, focusing instead on testing and installation. PROJECT EXAMPLES Hastings Bridge (Hastings, Minnesota) Dan Brown and Associates, PC (DBA) was the Foundation Engineer for the recently completed Hastings Bridge spanning the Mississippi River in Hastings, Minnesota (Dan Brown and Associates 2010). The key items in this case history include: (1) the increased reliability of the foundation design through demonstrated pile resistance; (2) the issues of vibrations on existing structures; (3) consideration of the limitations of dynamic tests to demonstrate fully mobilized pile resistance for piles driven to refusal on rock; (4) the use of a lateral load test for design; and (5) designing a test program for more than just “verification” of the design. The Hastings Bridge was constructed adjacent to an exist- ing bridge structure built in the 1950s. The old bridge was founded on large groups of 50-ft-long timber piles that were tipped primarily in fine-grained soils, but with medium-dense granular soils beneath the pile tips. The piles were driven as deep as practical for the era, particularly because they were being installed over water. To maintain the necessary navi- gation channel and to reduce the required span length the substructure locations for the new bridge were close to those for the old bridge. The new bridge includes five piers founded on groups of 42-in-diameter open-ended pipe piles. Information regarding the piers where LDOEPs were used is provided in Table 3 and Figure 15; Figures 16 and 17 show pile installation. The geology in this area of Minnesota includes some interesting foundation challenges, including up to a several hundred feet of highly organic and compressible very soft silts and clays to very dense sand and gravel overlying sedimentary bedrock. Experiences from five previous bridge projects using the same pile sections and constructed in the last 15 years in Minnesota and in similar geologic conditions suggests that these piles penetrate even dense granular materials easily to achieve bearing on rock. In all of these previous instances, a Delmag D125 diesel hammer (Rated Energy = 314,000 ft-lbs; Ram Weight = 27.6 kips) was used. The dynamic test results during initial drive and re-strikes did not fully capture the available geotechnical resistance under static loading condi- tions in either the soil or for piles bearing on rock. Initially, the potential impact of vibrations and vibration- induced settlements to the adjacent bridge during pile driv- ing for the new bridge was a significant concern. Some piles were located within 20 ft of the existing bridge and the instal- lation did produce some vibration that was noticeable, but nothing of any consequence and no damage whatsoever to the existing bridge, which was old and in poor condition. An automated and remotely operated instrumentation system was attached to the existing structure at several locations to continuously monitor displacement, tilt, and vibration and provide immediate alarm should any thresholds be exceeded. This system was augmented by manual optical survey mea- surements at specified time increments throughout the foun- dation construction duration. Based on these measurements, no adverse effects to the existing structure were observed. It can be postulated that open-ended pipe piles are probably advantageous in that regard because, despite a little bit of vibration associated with driving, there is no appreciable soil displacement as the pile cuts through the soil without plug- ging during installation.

39 TABLE 2 CASE HISTORY SUMMARY Name Location Type Pile Type Pile Diameter Hammer Type Soil Type Testing Method Notes Reference Hastings Bridge Minnesota Project Steel pipe 42 inch Diesel Rock Dynamic; Statnamic® Test method comparison; design Dan Brown and Associates (2010) Stoney Creek Bridge California Project Steel pipe 96 inch Diesel; hydraulic Dense sand, stiff clay, gravel Dynamic, Static Test method comparison; design Liebich (2009) Woodrow Wilson Bridge Virginia/ Maryland Project Steel pipe 54, 42, 36 inch Hydraulic Stiff clay; dense sand Dynamic; Static; Statnamic® Test method comparison; design Ellman (2009) St. George Island Bridge Florida Project; installation problems Concrete cylinder (spun- cast) 54 inch Not listed Florida limestone Dynamic; Static; Statnamic® Test method comparison; design; pile damage Kemp and Muchard (2007) Cross Bay Boulevard over North Channel New York Project Concrete cylinder (spun- cast) 54 inch Hydraulic Alluvium; sense sand Dynamic; Static Test method comparison; design verification NYSDOT (1996) Rigolets Pass Bridge Replacement Louisiana Project Concrete cylinder (spun- cast) 66 inch Hydraulic Clay Dynamic; Statnamic® Load test method comparison; design Robertson and Muchard (2007) Trout River Bridge Florida Installation problems Concrete cylinder (bed-cast) 54 inch Hydraulic Florida limestone Dynamic Pile damage Kemp and Muchard (2007) CAPWAP®- Based Correlations Alaska Research report Steel pipe 12 to 48 inch Diesel; hydraulic Dense sand; glacial deposits Dynamic Design method developed from CAPWAP® data Dickenson (2012) Kentucky Lake Bridge Kentucky Load test program Steel pipe 48 and 72 inch Hydraulic Chert residuum (Fort Payne) Dynamic; Static; Statnamic® Constrictor plates used to force plug Terracon (2014) Axial Pile Florida Research Steel pipe; 36 to 84 Diesel; Sand, clay, Dynamic; Design McVay (2004) Capacity of Prestressed Concrete Cylinder Piles report concrete cylinder (spun-cast and bed- cast) inch hydraulic Florida limestone Static; Statnamic® methods developed Oregon Inlet North Carolina Load test program Concrete cylinder (spun- cast) 66 inch Hydraulic Sand; silt Dynamic; Static Design Keaney and Batts (2007) Comparison of Dynamic and Static Tests Offshore Research report; load test program Steel pipe 36 to 78 inch Diesel; hydraulic Sand; clay; silt Dynamic; Static Test method comparison; design Stevens (2013) US-378 Bridge over Pee Dee River South Carolina Load test program Concrete cylinder (spun- cast) 54 inch Hydraulic Clayey sand; sandy silt (Pee Dee) Dynamic; Statnamic® Plug formation; test method comparison S&ME (2008) TABLE 3 HASTINGS BRIDGE LDOEP GROUP INFORMATION Pier Group Configuration Wall Thickness (in.) Length Prior to Cut-Off (ft) Embedded Length in Soil (ft) 6 3 x 7 Rectangle 1 175 135 7 2 x 7 Rectangle 1 185 165 8 2 x 7 Rectangle 1 185 170 9 4 x 2 x 4 (see Figure 15) 7/8 190 175 10 4 x 2 x 4 (see Figure 15) 7/8 185 175

40 The pile design was an iterative process between the foundation engineer and the structural engineer. The open collaboration was necessary to optimize the design by uti- lizing the full structural and geotechnical capacity avail- able. Considering the lateral demand in addition to the axial demand was necessary to optimize the pile section. Ultimately, the design was optimized to a point where both lateral and axial considerations produced controlling con- ditions under various load combinations. The 42-in.-diameter pipe pile was selected because it was considered to be the largest section that could be driven with the equipment available to the design-build contractor. Additional evaluations were performed considering a greater number of smaller piles; however, the larger diameter piles were more attractive given the efficiency in resisting lateral demand. In addition, the plan dimension size of the required pile cap and coffercell could be reduced by using fewer larger diameter piles. Upon arriving at the conclusion that 42-in.-open-ended pipe piles represented the preferred pile, the necessary wall thickness was determined. Accordingly, for the substruc- tures exposed to large vessel collision forces transverse to the bridge, the necessary wall thickness was determined to be FIGURE 15 Foundation plan view at Hastings Piers 9 and 10. FIGURE 17 Piers 6 and 7 under construction.FIGURE 16 Driving Pier 6 test pile.

41 1 in. At other locations not exposed to the fully loaded vessel forces, a 7/8-in. wall thickness was necessary as controlled by the strength limit state load combination in the longitudinal direction of the bridge. The pile wall thickness was deter- mined in order to resist the flexural demand, as well as the large axial demand on the outer piles resulting from group action under large lateral loading events. The amount of soil lost to scour was another important design consideration for this transportation structure. The large diameter pipe piles provided the necessary structural capacity in flexure to resist elevated bending moments follow- ing various scour events that had to be associated with corre- sponding load combinations. A lateral load test was performed on a test pile in the river to calibrate the soil model used for lateral load design. Finally, the potential for corrosion was evaluated in the design given the required 100-year design life specified by the owner. To account for the potential loss of section as a result of corrosion, a slightly larger wall thickness was specified than needed for structural resistance alone, as recommended by the project corrosion engineer to satisfy the minimum 100-year design life requirement in the Request for Proposal (RFP). Both dynamic and rapid (Statnamic®) load tests were per- formed for verification of the large axial resistance available for these piles bearing on rock (Dan Brown and Associates 2010). The implementation of Statnamic® load testing also allowed a higher resistance factor to be used in design at the Strength Limit State (when the factored load exactly equals the factored resistance) in accordance with the project RFP and at the discretion of the designer. Two axial Statnamic® tests were performed using a 10,000-kip gravel-catch device. One test was performed in the river on a pile with a 1 in. wall and the other on the north bank on a pile with a 7/8 in. wall. The results indicated approximately 4,600 kips and 4,200 kips of nominal resistance for the 1-in. and 7/8-in. wall piles, respectively. As anticipated, the behavior of the piles dur- ing the load testing was essentially elastic, as the pile heads deflected a maximum of about 2½ in. and rebounded almost completely with permanent sets of around ¼ in. The dynamic tests indicated about 3,000 to 3,500 kips of resistance with the piles bearing on rock at hammer blow counts indicating practical refusal. Therefore, by definition, that resistance is all that the hammer was capable of demonstrating and the nominal resistance used in design was that determined by Statnamic® testing. Dynamic testing was utilized on many of the production piles to demonstrate: (1) that the piles were driven to a good seating on rock; (2) that the piles were not damaged during installation; and (3) that the hammer was per- forming as intended. Stoney Creek Bridge (California) Caltrans published a case history for the Stoney Creek Bridge Project in “High-Capacity Piles at the Stony Creek Bridge Project” (Liebich 2009). This paper presents Caltrans’ expe- rience with the advantages and limitations of LDOEPs for a specific project. It presents a successful case of utilizing both dynamic and static testing to realize significant cost savings by using LDOEPs rather than smaller piles. The case also presents a comparison of field tests with design calculations that were based on unplugged behavior under static loading: a portion of the side resistance on the interior of the pile was included in the geotechnical resistance. As described in this paper, Caltrans uses the terminology “high-capacity piles (HCPs),” which are defined as piles that are larger than 3 ft in diameter or have an axial capacity of greater than 2,000 kips. The piles for the Stoney Creek Bridge Project are 8 ft in diameter, have a wall thickness of 1.75 in., and are embedded 170 ft below grade. Stoney Creek Bridge is located approximately 100 miles north of Sacramento on Highway 32. The original bridge was completed in 1976, but by 1992 unanticipated scour conditions had partially exposed the bridge foundations. A significant ret- rofit was performed, including structural elements to protect bridge piers and extensive placement of rip-rap. By 2000 how- ever the structure was found to again require remediation. It was then determined that, as a result of nearby mining opera- tions, the potential scour depth could eventually reach 50 ft. For the new design, a single LDOEP was located under each of the seven single-column bents. The required pile nominal geotechnical resistance was calculated as 7,800 kips utilizing the full external side resistance along the perimeter of the pile, the base resistance using the wall area of the pile, and the internal side resistance gained from the soil in the lower third of the pile. The soil profile generally consisted of layers of sand, clayey sand, clayey gravel and fat clay with hard to very stiff fine sandy clay, or sandy silt to below expected tip. The first production pile was installed as a test pile for a static axial load test. To verify the pile resistance for the maxi- mum design scour condition, the test pile was installed within a sheet pile cofferdam excavated to the approximate scour elevation to isolate the pile from side resistance in the scour zone. The pile was initially driven using an APE D100-13 diesel hammer (Rated Energy = 248,000 ft-lbs; Ram Weight = 22 kips), but was completed using an APE HI 750U hydrau- lic hammer (Rated Energy = 750,000 ft-lbs; Ram Weight = 120 kips), owing to high penetration resistance encountered by the smaller hammer. Liebich (1999) reported that wave equation analyses before construction indicated that the larger hammer would not be sufficient to demonstrate the required resistance after pile setup; therefore, a static test load was planned. A static axial load test was conducted 28 days after instal- lation, attaining a resistance of 7,860 kips before experi- encing a plunging geotechnical failure (Figure 18). Liebich notes in the paper that the load test matching so well with the calculated resistance is more the result of favorable coinci- dence than repeatable skill.

42 Woodrow Wilson Bridge (Virginia/Maryland) The replacement of the Woodrow Wilson Bridge between Alexandria, Virginia, and Prince George’s County, Mary- land, was completed in 2009. A major test pile program was executed to enable efficient design of the planned driven pile foundations (Ellman 2009). The program included load tests on three 54-in.-diameter, three 42-in.-diameter, and one 36-in.-diameter steel pipe piles. All piles had 1-in.-thick walls and all were installed with a ICH S-280 (Rated Energy = 206,000 ft-lbs; Ram Weight = 30 kips) or a ICE-275 (Rated Energy = 110,000 ft-lbs; Ram Weight = 27.5 kips) hydraulic hammer. Testing included PDA® monitoring during driving as well as 7-day and/or 14-day restrikes. Signal matching analyses using CAPWAP® were performed for all dynamic tests. Static load tests were performed on one of each of the three pile sizes. Statnamic® rapid load tests were performed on one 54 in. and one 42 in. pile. The soil conditions at the bridge site included very deep soft alluvium along much of the alignment, especially at the area of the main channel where the new bascule span was to be located. For the bascule span, the piles would be driven through the alluvium to bear in a stiff to hard clay. Parts of the approaches would bear in the same clay, while others would terminate piles in a dense sand stratum. Measurements of the side resistance by dynamic testing during driving indicated to the design team that the piles did not plug during driving. Side resistance values computed from CAPWAP® analyses were relatively uniform and low (as expected) for the soft alluvium. The CAPWAP® calcu- lated side resistance in the underlying clay was much more variable than expected. The data from all of the test piles was used to develop an average profile from which select val- ues of side resistance were selected for each stratum. These values were used in wave equation analyses to predict blow counts during driving for comparison with the observed pile driving. However, as noted later, the observed blow count ended up not being the pile acceptance criteria. The static and Statnamic® load test piles were instrumented to measure the distribution of side resistance and the propor- tion of side versus base resistance. The total axial resistance and unit side resistance measured by the Statnamic® tests were significantly higher than the values measured by the static tests Compressive Load (kN) <1 kN = 4.448 kips> FIGURE 18 Load-displacement plot from Stoney Creek Bridge load test (Liebich 2009).

43 as shown in Table 4. Because of the significant differences, the project team decided that there was poor correlation between the two test methods at this site; therefore, the team conserva- tively used the lower static test result values. Ellman noted that the team believed the static test data indicated the pile did behave in a plugged manner under static loading. He also noted that unit side resistance values in the stiff to hard clays appeared low (0.9 to 2.3 ksf) based on the relatively high (7 to 8 ksf) undrained shear strength of the clays. Published test data would indicate values of 3 to 4 ksf for similar strength clays. Ultimately a design value of 1.7 ksf for unit side resistance was selected based on all of the test data and engineering judgment of the designers. Back-calculation of alpha values from the selected unit side resistance values was not included in the paper. After analysis of all of the test data and performing driv- ability studies, the design team decided to use a “Specified Tip Criteria” for production pile installation. These criteria required piles be installed to a set tip, not a minimum dem- onstrated resistance. The relative uniformity of the subsur- face conditions for the water spans led the team to select this method to simplify pile installation. The criteria still included dynamic testing to perform quality assurance/quality control testing and check driving stresses. The design phase testing was deemed to be successful in allowing the optimization of the foundations for the bridge. The results of the test data were extrapolated to design 72-in.- diameter open-ended pipe piles for the bascule span foun- dations. The Maryland approach for foundations consisted of 48-in., 54-in., and 66-in.-diameter piles (the Virginia approaches were designed for 24-in.-square precast concrete piles). The use of the Specified Tip Criteria was reported to have worked well for the pipe pile installations. St. George Island Bridge (Apalachicola, Florida) This project utilized segmented spun-cast post-tensioned con- crete piles for a new bridge over Apalachicola Bay (Kemp and Muchard 2007). Comparisons of static, rapid (Statnamic®), and dynamic tests were performed on four test piles that included embedded strain instrumentation. FDOT uses standard pile designs for segmented spun-cast post-tensioned concrete piles (see FDOT interview notes). One of the standard designs consists of a 54-in. outside diameter pile with 8-in.-thick walls, and a specified concrete strength of at least 6,000 psi at the time of application of the pre-stressing force. A 28-day compressive strength of 7,000 psi was used for this particular project. The advantage of using spun-cast cylinder piles for the project was that it allowed the casting yard to start production of pile segments while still awaiting the test pile results to determine the final order lengths of the post-tensioned assembled piles. Casting segments early in the project helped lessen material escala- tion costs as well as pile driving rig down time for the 646 piles planned for the bridge. The subsurface conditions at the site generally consisted of fine silty sands over weathered Florida limestone (also called “limerock”). The top of the limestone into which the piles were embedded varied in elevation from -60 ft to -80 ft National Geodetic Vertical Datum (NGVD). The factored load for the piles ranged from 600 to 1944 kips. The load test program included four static load tests, six Statnamic® load tests, and 50 dynamic tests on production piles. Figure 19 shows the static and Statnamic® test setups. The static and Statnamic® test piles were also subjected to dynamic tests, including CAPWAP® analyses of restrikes. Reported test results indicated that the static and Statnamic® tests were in reasonable agreement, with the Statnamic® resis- tance reporting 2% less than the static test resistance for three of the four test piles, and 9% less for the fourth pile. It was also reported that the pile resistance estimated by signal match (CAPWAP®) on the load test piles was 9% to 42% less than the static test resistance, although the time between re-strike and static test, and the details of the signal match models (e.g., plug, radiation damping) were not specified. Table 5 shows the total resistance values. Cross Bay Boulevard over North Channel (Queens County, New York) NYSDOT compiled a case history of the test pile program for a replacement bridge for Cross Bay Boulevard over the northerly section of Jamaica Bay in Queens County (NYSDOT 1996). The purpose of the case history included an evaluation and cor- relation of the static and dynamic pile load tests performed on the project in order to develop recommendations for projects utilizing similar foundations in similar soil conditions. Location (Hammer) and Pile No. Pile No. Pile Diameter (in.) Total CAPWAP® Resistance (kips) Static Test Resistance (kips) Statnamic® Test Resistance (kips) PL-1 (IHC S-280) B 54 4250 5300 PL-1 (IHC S-280) C 54 3133 2929 PL-2 (IHC S-280) E 42 2948 4360 PL-2 (IHC S-280) F 42 2786 >2000 PL-3 (ICE-275) I 36 1324 >1900 TABLE 4 WOODROW WILSON BRIDGE TEST PILE RESULTS

44 Construction of the replacement bridge structure began in 1988. The northbound half of the bridge was completed first and opened to traffic in 1991. This allowed for the demolition of the existing bridge and construction of the southbound half, which was completed in 1993. The six-lane bridge is 2,842 ft long, supported on two abutments and 33 piers. The approach spans are supported on pile bents, each with six 54-in.-cylindrical spuncast concrete piles. The main channel span Piers 7 and 8 are each supported on sixty 14-in.-square piles. The maximum reported design load on the 54-in. cyl- inder piles estimated during design was 660 kips. Only air or steam hammers were allowed to drive the piles. According to the report, “The bridge is located on the south shore of Long Island, an area where the glacial outwash soils (fine to coarse sands) from the Long Island terminal moraine have been modified by wave and wind action. The density of the sands ranges from very loose near the surface to dense and very dense below elevation -65 ft to -75 ft. Soft organic clayey silt, about five to ten feet thick, deposited during post- glacial times in the sheltered waters of Jamaica Bay, covers the sandy soils in the North Channel. The channel is connected to the open sea and is subject to tidal water level fluctuations.” Pile 3 of Pier 26 was subjected to dynamic and static testing. The test pile was jetted to elevation -54 ft and then driven with a CONMACO 5300 hydraulic hammer (Rated Energy = 150,000 ft-lbs; Ram Weight = 30 kips), using a 3 ft stroke, to elevation –75 ft. The pile was then driven an additional 6 in. using a 5 ft stroke. Dynamic testing with FIGURE 19 Static (left) and Statnamic® (right) load tests at St. George Island Bridge [Photos by Applied Foundation Testing (Copyright 2000)]. Source: Kemp and Muchard (2007). Pile No. LT-1 LT-2 LT-3 LT-5 CAPWAP® Restrike Maximum Capacity (kN/tons) 8,667/975 8,960/1,008 8,089/910 9,013/1,014 Static Load Test Maximum Capacity (kN/tons) 9,493/1,068 13,813/1,554 13,600/1,530 12,836/1,444 Statnamic® Load Test Maximum Capacity (kN/tons) 9,627/1,083 13,564/1,526 13,831/1,556 11,689/1,315 TABLE 5 SUMMARY OF TEST RESULTS FOR ST. GEORGE ISLAND BRIDGE

45 CAPWAP® analysis indicated a resistance of 659 kips at the EOID. Based on these test results, driving criteria were established that incorporated the jetting as well as both the 3 ft and 5 ft stroke settings. After completion of the restrike test, a static load test was performed on the pile. The “failure load,” or maximum resis- tance, as interpreted by NYSDOT was 1548 kips, 2.35 times the estimated resistance of 660 kips. The piles did not include any embedded strain gauges to measure the distribution of side and base resistance. To further investigate the pile behavior, NYSDOT per- formed a more detailed CAPWAP® analysis after the com- pletion of the project. For this analysis, NYSDOT made two different assumptions regarding the pile plugging behavior: (1) The pile and plug act separately (pile “cookie cuts” through the soil); and (2) the soil plug in the pile is driven together with the pile and adds to the mass of the pile. Pile resistance was calculated separately for each assumption. From these analyses, NYSDOT determined that the detailed CAPWAP® analyses showed relatively good agreement with the static load test at pile displacements of about 0.4 in. This was not initially the case immediately at the conclusion of the test program, when the dynamic testing appeared to be under- estimating the total pile resistance. Only after evaluating the data and accounting for plug behavior did NYSDOT develop an approach that provided good agreement between the dynamic and static test results. Although no direct correlation between dynamic testing and anticipated static pile resistance was developed, NYSDOT applied the knowledge gained to future projects in the form of better understanding the level to which dynamic testing can under-predict pile resistance in similar conditions. Rigolets Pass Bridge Replacement (Slidell, Louisiana) This project was for the Rigolets Pass Bridge on US-90 near Slidell, Louisiana (Robertson and Muchard 2007). Statnamic® load testing was performed on three 66-in.-diameter, 6-in.- thick wall spun-cast, post-tensioned, concrete cylinder piles. The case history focused on comparison of the Statnamic® results with the dynamic testing. The soil conditions consisted of sands overlying intermit- tent layers of dense fine sand and hard clay transitioning to dense fine sand, with the piles tipped into the hard clay layer. The piles were cast with strain instrumentation at four loca- tions along the length of the pile. A Bruce SGH-3013 hydrau- lic hammer was used to install the test piles (Rated Energy = 282,000 ft-lbs; Ram Weight = 66 kips). Dynamic testing with PDA® was performed on all three test piles during installa- tion, including restrikes. The pile resistance values calculated from the Statnamic® tests and the CAPWAP® analysis of the re-strikes are shown in Table 6. Analysis of the test data suggested that the dynamic testing did not fully mobilize the side shear resistance as well as the Statnamic® testing. Their conclusions concerning the differences in CAPWAP® and Statnamic® resistance values for this project were the result of (1) the hammer may have had sufficient energy to install the piles but not sufficient energy to fully mobilize the nominal static resistance during re-strikes, and (2) soil plugging behavior that is present dur- ing Statnamic® testing may not exist during dynamic testing. REPORTED INSTALLATION PROBLEMS AND ADOPTED SOLUTIONS St. George Island Bridge (Apalachicola, Florida) In addition to the load testing, this case history (Kemp and Muchard 2007) also highlights the investigation and mitiga- tion of longitudinal cracking that occurred in some of the piles during driving. About two-thirds of the way into pro- duction pile installation longitudinal cracks were observed in some piles, usually within three to four weeks after driving. The cracks were noticed primarily from evidence of calcium efflorescence (or leaching) that developed. Eventually 7% of the driven piles were determined to have cracks. Observation and mapping of the cracks was performed to further evaluate the problem. The investigation of the cracking included a review of the pile structural design and the performance of the driving sys- tem. Neither of these was determined to be out of the ordi- nary. Searching available engineering literature on the subject revealed that the problem was not unique, but that very little had been done to isolate the cause or causes. The paper ref- erences two significant resources: a report by the Louisiana Transportation Research Center (Avent and Mukai 1998) and a report by the PCI Journal (PCI 1993). The Avent and Mukai report (1998) investigated cracks post-construction; therefore, it did not have specific conclusions as to the causes Test Pile Pile Diameter (in.) Total CAPWAP® Resistance (kips) Statnamic® Test Resistance (kips) TP-2 66 1245 2966 TP-3 66 2030 3077 TP-4 66 2166 3315 TABLE 6 RIGOLETS PASS BRIDGE TEST PILE RESULTS

46 of the cracks. The report did note that no significant corrosion had occurred, even at ages of up to 40 years. The PCI reported that (1993) “states when driving piles in semi-fluid soils, a soil plug may come up through the hollow of the pile building up a ‘hydraulic ram effect’ and produces high circumferential tensile forces exerted on the pile walls. This internal loading condition occurs only during installation. It is also known as ‘water hammer’ and results in what is sometimes referred to as ‘hoop stress.’ These little understood forces apparently can exceed the pile’s hoop stress resistance causing the cracks.” Solutions to reduce the potential for these forces to build up include increasing the number and size of vent holes in the piles, increasing the lateral reinforcement of the pile, or peri- odically cleaning out the plug. Since the pile segments were already cast for the project the contractor elected to clean out the plug. This resulted in no further cracking observed during installation of the remaining 250 piles on the project. Trout River Bridge (Jacksonville, Florida) The Trout River Bridge project (Kemp and Muchard 2007) utilized 54-in. bed-cast concrete piles. The case history focused on the issue of severe spalling of the pile top during instal- lation. Dynamic testing was utilized to monitor pile stress and confirm nominal pile resistance, but no static tests were performed. The factored design load for the piles was 550 kips. To achieve the required resistance the contractor opted to use an APE 400 hydraulic hammer (Rated Energy = 320,000 ft-lbs; Ram Weight = 80 kips). The rated energy exceeded the mini- mum requirement of 240,000 ft-lbs included in the project plans. For this project, FDOT used a pile reinforcement arrangement that had an increase in spiral ties throughout the pile length, compared with previous designs. Despite the increased reinforcement in the piles, the first three piles experienced substantial cracking and spalling problems near the top three to five feet of the pile. The first pile had such sever spalling that the pile head had to be cut off below the damaged portion in order to continue driving. The investigation of the root cause of the cracking even- tually focused on the driving system and the geometry of the custom-made helmet, which was suspected of causing local- ized large driving stresses. The particular hammer used by the contractor had been special ordered for the project and included a driving bell to maintain alignment of the pile within the driving system. The driving bell had an inner ring to hold the pile cushion in place above the pile. This ring extended a relatively small distance into the pile void when the ham- mer was set on the pile. The driving bell also had ribs along its skirt in four places that fit closely to the outside of the pile. Up to ½ in. of free movement of the pile was possible between the inner and outer restraints. The free movement allowed the mechanism to be out of vertical alignment, caus- ing the pile to get “pinched” in the bell during each blow of the hammer. The pinching action from the tilted bell resulted in cracks that eventually led to spalling after sufficient ham- mer blows. The reported solution was to add shims between the inner and outer alignment ribs within the driving bell (Figure 20) to reduce the free play that had developed. After adding wedge plates as shims and removing the inner diameter pilot ring the extent of cracking and spalling was greatly reduced. Only infre- quent minor cracking was observed after the modifications. LOAD TEST CASE HISTORIES AND RESEARCH REPORTS During the literature review for this report several published case histories highlighting load test programs on LDOEPs were found, as well as research reports investigating pile behavior. The following is a list of some of the notable case histories with a brief summary of the project conditions and pile configurations. These are presented to allow the reader to quickly identify published histories or reports that may be useful for a specific project or condition. CAPWAP®-Based Correlations for Estimating the Static Axial Capacity of Open-Ended Steel Pipe Piles in Alaska This report (Dickenson 2012) compiled CAPWAP® analy- ses of data from PDA® monitoring of open-ended steel pipe piles at 32 bridge sites across Alaska. The piles ranged in diameter from 12 in. to 48 in., with depths ranging from 23 ft to 161 ft in mostly cohesionless soils. The database contains analyses for 68 piles, 33 of which include data for EOID and beginning of restrike. The report presents region-specific, empirical relation ships for unit shaft resistance and unit toe resistance at BOR. FIGURE 20 Inside driving bell (after Kemp and Muchard 2007).

47 The predominate soil type for bridge projects in Alaska, and hence the predominate soil type in the database, is cohe- sionless soil deposits. The deposits in the state can be broadly categorized as Class 1 (normally consolidated) and Class 2 (highly overconsolidated). Class 1 is the most common, with Class 2 being found primarily in the Anchorage area. For esti- mating unit side resistance the report provides a CAPWAP®- based relationship of stress-normalized unit shaft resistance (fs/sv′) with depth developed from the database. The relation- ship for Class 1 soils is shown in Figure 21. For unit side resistance in cohesive soils, the report recommends the alpha and beta methods in the current FHWA design manual (i.e., Hannigan et al. 2006). For unit toe resistance, the report recommends the meth- ods outlined in the FHWA manual supplemented with the range of “equivalent” unit toe resistance values presented in the report computed from CAPWAP® analyses in the database. Dickenson designates the toe resistance values as “equivalent” owing to the impact on how the PDA® data are evaluated with respect to assumptions of pile plugging. The proposed relationships were used to predict pile resis- tance on a project with 29 monitored piles driven into silt-rich deltaic deposits. Dickenson reported “Overall, the agreement between the predictions and the CAPWAP® results was good to excellent, and the proposed method provided much more reliable ranges of estimated pile resistance than obtained using widely adopted, standard of practice procedures.” Kentucky Lake Bridge Pile Load Test Program A design-phase pile load test program was performed to study the constructability and design of 48-in. and 72-in. steel LDOEPs for the Kentucky Lakes Bridge project in Marshall and Trigg counties, Kentucky (Terracon 2014). The piles were installed with a Menck MHU 800S hydraulic hammer (Rated Energy = 604,000 ft-lbs; Ram Weight = 100 kips). The testing included dynamic, static axial, Statnamic® axial, and Statnamic® lateral tests on piles with wall thicknesses ranging from 1 to 2 in. Vibration monitoring was also performed. There were several design issues that led to consider- ing LDOEPs and having a test pile program for this par- ticular project. The two main factors were the efficiency of fewer, larger piles providing greater stiffness for seismic and impact loads, and the difficult soil and/or geologic conditions for drilled shafts and smaller piles at the site. The site was underlain by residuum of the Fort Payne Formation—a chert residuum that behaves as a dense gravel, with some silt and clay layers. This formation creates potential difficulties with keeping drilled holes open, usually requiring full-depth tem- porary casing. The formation is known to create difficulties in achieving clean shaft bottom conditions and is also known to cause excessive wear on drilling tools, driving up costs for contractors. Other pile types were problematic because closed-end piles would not be able to reach the bearing strata and there was concern that small diameter open-ended piles would plug easily and not reach the bearing strata. An important question for the project was how could the certainty of the plug developing and its location be bet- ter determined? The location of a significant chert stratum directed the design details in that the plug would be relied on to either achieve penetration into the chert or to achieve the nominal pile resistance. The team wanted to keep the pile open to get it down relatively easily, but needed it to plug to get the pile to drive the minimum distance into the chert required for fixity. It was believed that having the piles plug and thus act as a displacement pile when penetrating the chert would allow the piles to more easily penetrate the chert than if the piles cut into it. There was concern on the part of the design team that the pile toe would be damaged if the piles attempted to cut into the chert rather than displace the material. The team used the load test program to experiment with where to create the plug, designing steel constrictor plates to be installed inside the pile to attempt to form the plug at a fixed location. A sche- matic of the one of the plates is shown in Figure 22. The design evaluations included discussions between the design team and Paikowsky and his work on the Sakonnet River Bridge in Rhode Island (Paikowsky 2011). Another aspect of the testing program was a detailed evaluation of dynamic pile testing records to consider meth- ods of improving the match quality and the estimation of static pile resistance when considering plugging. Comparisons FIGURE 21 Trend of stress-normalized unit shaft resistance (fs/sv´) with depth for site Class 1 soils (Dickenson 2012).

48 of analyses using a single-toe pile model and a double-toe model were made, as well as application of radial or radiation damping models. The radiation model assumes some energy is radiated away from the pile tip instead of being completely confined to static and dynamic responses of the soil shear along the pile and at the pile toe. The following concerning the analyses can be reported: • For the dynamic records where radiation damping was applied the model generally resulted in a significantly better signal match quality, indicating that the radiation damping allows CAPWAP® to better model the signals recorded by the dynamic pile testing equipment. • The pile resistances calculated with CAPWAP® using the radiation damping model also generally produced higher end-bearing resistance values than the CAPWAP® models without the radiation damping. It appears that the radiation damping model is better suited for estimating the end-bearing component of the piles when less pile set is experienced per hammer blow. This is the case when the constrictor plates are engaged on the dense granular soils. • Wave equation analyses indicated that plugged piles would have high stresses. In addition, there was con- cern that localized high stresses might be encountered owing to the presence of the chert. Testing on the piles typically did not approach as high values as expected. Terracon and KYTC used the API RP 2A method for their comparative analysis of static resistance calculations to the load test results. The experience of Terracon with this method for similar piles is that it provides a better prediction of static resistance than the FHWA methods. An example of their analyses is provided in Figure 23, which shows the load test and calculated static resistance for the 48-in.-diameter piles in the shallow water test location. Table 7 lists the reported CAPWAP® and estimated static resistance of the test piles from the static or Statnamic® tests. Some of the key conclu- sions reported for the testing program were: • The tests indicated that the 48-in. piles were more likely to achieve plugging with the plates at the target stratum for the load test program than were the 72-in. piles. The results of the test program were used to revise the target stratum for positioning the plates for the production piles. • Vibrations generated by driving the selected piles would not likely be damaging to the adjacent existing bridge; however, monitoring was recommended for all adjacent existing bridge piers during production pile installation. • Dynamic testing was generally underestimating the static pile resistance as determined by the static and Statnamic® tests. The use of radiation damping models generally reduced the magnitude of the underestimates. The driv- ing criteria for production piles can be developed based on the lower dynamic resistance correlated to the static resistance. Determination of Axial Pile Capacity of Prestressed Concrete Cylinder Piles This research (McVay 2004) evaluated steel and concrete LDOEPs under driving and static loading conditions, includ- ing special attention to plug behavior under both conditions. Load test data for 35 tests (22 concrete piles and 13 steel piles) on pile diameters ranging from 36 to 84 in. from 11 projects were assembled and analyzed. The data were obtained from FIGURE 22 Constrictor plate schematic (courtesy: Genesis Structures, Inc. and KYTC).

49 FDOT, Caltrans, VDOT, NCDOT, and MSHA. A summary of the work was also presented in Lai et al. (2008). The load test analyses focused on developing equations for unit side resistance and unit end bearing as a function of SPT N values. Two approaches were used depending on the data available for each test: a Direct Method using strain gauge data from instrumented piles and an Indirect Method plotting load versus deflection data on arithmetic scale to determine the Davisson’s Capacity (“failure” load) and on a log–log scale to use DeBeer’s Method to determine the distribution of side and tip resistance. The report provided relationships for unit side and base resistance as a function of SPT N for different soil types Analysis of pile plug behavior during driving was evalu- ated by a parametric study of pile diameter on inertia force and required soil column height inside the pile to generate enough friction to cause plugging to occur. This study was done for one soil type (silt) and one pile length (80 ft) for inside pile diameters of 54, 38, 24, 20, 18, and 12 in. For each diameter a critical g-force was determined for the soil column. For pile accelerations below this level, the soil plugs the pile and the soil column moves down with the pile. For pile accelerations above the critical g-force the pile does not plug and “cookie cuts” into the soil. In general, the report noted that prestressed concrete piles driven in accordance to FDOT specifications typically have accelerations from 40 g to 60 g, well above the critical g-force values for typical concrete pile diameters (15 g for a 54-in.-diameter pile). To evaluate potential plugged behavior under static load- ing, finite element modeling was performed. The analysis suggested that for typical pile wall thicknesses, soil strengths, FIGURE 23 Load test and calculated geotechnical resistance analysis (Terracon 2014). Test Pile Pile Diameter (in.) EOD CAPWAP® Resistance (kips) 72-hour Restrike CAPWAP® Resistance (kips) Estimated Static Resistance (kips) K-1 48 3300 5000 9550 (Static test) K-2 48 3250 4730 6952 (Statnamic® test) K-3 72 3800 5200 8511 (Statnamic® test) TABLE 7 KENTUCKY LAKES BRIDGE TEST PILE RESULTS

50 and rate of loading for FDOT projects, plugged behavior dur- ing static loading was feasible. LRFD resistance factors for design of concrete cylinder piles were also developed during this research. McVay used the advanced first order second moment method to perform the reliability calibrations and derive recommended resis- tance factors from the load test data. Two approaches were investigated; Approach 1 utilized only the “ring area” of the pile in the resistance calculations—only the cross section of the pile material was used, neglecting the void in the center. Approach 2 utilized the full cross-section area of the pile. Using a target reliability index, bT = 2.75 based on previous FDOT LRFD calibration work; a resistance factor of j = 0.76 was obtained for Approach 1 and j = 0.61 for Approach 2. Oregon Inlet (North Carolina) In 1996, a 66-in.-diameter, 140-ft-long prestressed post- tensioned concrete cylinder test pile was driven and load tested at Oregon Inlet, North Carolina, for the design of a new replacement bridge (Keaney and Batts 2007). The pile was driven using an HPSI 3505 hydraulic hammer (Rated Energy = 176,000 ft-lbs; Ram Weight = 35 kips). The pile was jetted and driven while monitored with PDA®, including a restrike test 16 hours after the EOID. Difficulties with maintaining hammer alignment were observed with the eccentric stresses measured by the PDA® during driving, but were deemed to not affect the pile penetration. At the EOID, the pile tip was stopped about 10 ft short of the planned tip elevation. During the restrike test, no addi- tional movement of the pile was observed, even with the jets engaged at full pressure. The contractor ended up cutting off the top 10 ft of the pile to accommodate the designed reaction frame system for the static axial load test. The static load test was performed to a maximum test load of 1962 kips (220% of the pile design load). The authors reported a resistance of 1266 kips calculated from CAPWAP® analysis of the restrike data, but noted that the restrike did not fully mobilize the pile resistance. There was no mention of an attempt to use superpositioning of EOID and restrike data to calculate the pile resistance. The static test resistance was about 50% greater than the CAPWAP® calculated resis- tance, indicating that the restrike did not fully mobilize the pile resistance. A Comparison of Dynamic and Static Pile Test Results A paper presented at the 2013 Offshore Technology Con- ference compared dynamic and static load tests performed on three offshore projects in mixed sand and clay profiles (Stevens 2013). The LDOEPs installed were steel pipe ranging from 36 in. (914 mm) to 78.7 in. (2,000 mm) in diameter. The observed and predicted pile resistance values were in reasonable agreement, including use of CAPWAP® analyses to evaluate contributions of side resistance. The first project included two 48-in.-diameter steel pipe piles driven with a Junttan HHK 14S hydraulic hammer (Rated Energy = 152,000 ft-lbs; Ram Weight = 31 kips) up to 65 ft into a sandy soil profile. Restrikes were performed 2 and 27 days after initial drive. CAPWAP® analyses of both initial drive and restrike showed little change in the shaft resistance and end bearing. The shaft resistance obtained from the CAPWAP® analyses was 1290 and 2530 kips at 40 and 65 ft penetration, respectively. These values were compared directly with the results of pull-out tests performed 52 and 46 days after restrike. The piles were reported to have a tension resistance of 1180 and 2530 kips at 40 and 65 ft penetration, respectively. The test results were considered to be in good agreement with the resistances predicted from the CAPWAP® analysis. The second project utilized 36-in.-diameter steel pipe pile initially vibrated with an HPSI Model 1600 vibratory driver and then impact-driven using a Berminghammer B-6505 diesel hammer (Rated Energy = 203,000 ft-lbs; Ram Weight = 17.6 kips) in dense to very dense silty sand to sand, stiff to very stiff clay, and very dense silty sand. The pile was not monitored during initial drive. CAPWAP® analyses were performed at the beginning and end of restrike to estimate the pile resistance as 4337 kips. A static load test was per- formed 5.5 months after driving the pile; however, it was only tested to 3400 kips (twice the design load of 1700 kips) with a maximum displacement of 1.5 in. The test was not conducted past the specified maximum test load in order to compare with the predicted CAPWAP® resistance. The third project included 78-in.-diameter steel pipe pile driven with a Menck MRBS 5000 steam hammer (Rated Energy = 542,000 ft-lbs; Ram Weight = 110 kips) to a pen- etration of 96 ft in interlayered very stiff clay and very dense sand. The test program was designed to evaluate pile setup. CAPWAP® analysis of PDA® data for the EOID and the beginning of restrike from one of the reaction piles was used to predict the test pile pull-out resistance for a static load test performed 52 days after driving. A delay of 43 hours occurred between EOID and restrike. An estimated dissipa- tion of pore pressures was used to calculate the potential setup 52 days after driving from the 2-day setup determined from the CAPWAP® analysis. A compressive static load test was performed with a test pile instrumented with strain gauges to determine the load distribution along the pile. The shaft resistance calculated from the load test was 5875 kips. This compared favorably with the authors predicted 5950 kips from the CAPWAP® data and pore pressure dissipation model. US 378 Bridge over Pee Dee River (South Carolina) The South Carolina DOT executed a comprehensive founda- tion test program for this bridge replacement project (S&ME 2008). The program included dynamic and Statnamic® load

51 testing of two 54-in.-diameter prestressed concrete cylin- der piles, an 18-in.-square prestressed concrete pile, and a 5-ft-diameter drilled shaft. The two 54-in.-cylinder piles (designated Cylinder Pile A and Cylinder Pile B) included embedded instrumentation. Both were installed with an APE 400u hydraulic hammer (Rated Energy = 400,000 ft- lbs; Ram Weight = 80 kips). The piles were installed through sandy overburden into the Pee Dee Formation, a calcareous clayey sand to sandy silt. To monitor the formation of a plug in the interior of the pile, a simple device called a pile plug monitoring device (PPMD) was constructed. The PPMD consisted of lead weights attached to a 100-ft fiberglass measuring tape. The weights would fall to the top of the soil column inside of the piles, allowing the distance to the soil to be computed. Access to the interior of the pile was made through a vent hole near the top of the pile. The PPMDs were read intermit- tently throughout test pile installation. The data showed that soil was rising inside both piles during driving indicating that disturbed soil and water was accumulating in the pile rather than a pile plug forming and traveling down with the pile. Dynamic testing of the cylinder piles was performed dur- ing installation and restrikes, followed by Statnamic® testing. Cylinder Pile A was dynamically monitored during a restrike 3 days after installation and again 20 days after installation (1 day after Statnamic® testing). Cylinder Pile B was only monitored during a 21-day restrike (1 day after Statnamic® testing). CAPWAP® analyses were performed on all dynamic tests. Statnamic® tests were performed on both test piles using a maximum derived static load of 2800 kips for Cylinder Pile A and 2500 kips for Cylinder Pile B. The strain gauge instru- mentation was used to calculate the load distribution and the unit resistance values for each pile. Figure 24 was presented in the report to summarize the dynamic and Statnamic® test results, including the calcu- lated unit side and base resistance values. It was determined that for Pile A the dynamic resistance was 20% lower than the resistance from the Statnamic® test, while the dynamic resistance was 10% higher than the Statnamic® resistance for Pile B. FIGURE 24 Test results summary for Pee Dee River Bridge load test (S&ME 2008).