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8 chapter two STATE OF PRACTICE AND LITERATURE REVIEW INTRODUCTION LDOEPs are prefabricated tubular steel or prestressed con- crete cylinder piles that are 36 in. outside diameter or larger and are driven into the subsurface to provide axial and lateral foundation support for the structure. These piles present a unique challenge for foundation designers owing to the com- bination of several factors: ⢠The tendency of the piles to âplugâ during installation is uncertain and may affect the behavior during installation, ⢠The potential for installation difficulties and pile dam- age during driving is unlike other types of conventional bearing piles, ⢠The soil plug within the pile may behave differently during driving or dynamic testing compared with static loading, ⢠Axial resistance from internal friction, and ⢠The nominal axial resistance may be very large and there- fore verification with conventional load testing is more challenging and expensive. This chapter provides an overview of the state of the prac- tice with respect to the use of these piles for transportation structures, including a review of published literature on the subject. SELECTION OF LARGE DIAMETER OPEN-ENDED PILES (LDOEPs) FOR TRANSPORTATION STRUCTURES LDOEPs have primarily been used for bridge structures where one or more of the following conditions exist: ⢠Lateral load demands on the foundation are relatively high, often as a result of extreme event loading condi- tions such as vessel collision or seismically induced lat- eral forces. ⢠The piles are subject to a significant unsupported length as a result of scour, liquefaction, or marine conditions. ⢠Soils are relatively weak to a fairly substantial depth. ⢠Axial demand on the foundation is high. ⢠The use of a single large diameter pile can eliminate the need for a footing, such as to allow the use of a pile bent for a pier substructure. ⢠Marine construction conditions are implemented for pile delivery, handling, and installation. LDOEPs are particularly favorable where large lateral demands must be resisted because of the significant flexural strength that is efficiently provided by a large diameter cylin- drical shape formed of high-strength engineered materials. In addition, these piles provide the advantages of ductility where seismic stresses may be high. Bridges exposed to deep scour can result in long unsupported pile lengths and high bend- ing stresses. Liquefaction conditions associated with seismic events increase flexural strength demand on piling. Vessel col- lision forces and other extreme event loadings can demand large lateral resistance from foundations and therefore favor the use of LDOEPs. It is not unusual for LDOEPs and drilled shafts to be compared as alternatives in many such cases (S&ME 2008), because many of the conditions cited previously also favor the use of drilled shaft foundations. The factors that most favor LDOEPs over drilled shafts are the presence of deep weak soils and/or marine construction conditions. Drilled shafts are most cost-effective where a strong bearing stratum exists that can be engaged to provide resistance. Conditions where lateral resistance requires embedment into a hard stratum such as rock are not favorable for driven LDOEPs; however, if the rock is deep and lateral resistance is provided by over- burdened soils, LDOEPs offer a potentially simpler and faster method of constructing deep foundations. The use of prestressed concrete cylinder piles for transporta- tion structures has been concentrated along the Gulf and Atlan- tic coasts, with cylinder pile sections of 36 in. to 66 in. outside diameter driven to bear in coastal alluvial sediments. A typical application for prestressed concrete LDOEPs might be to use these piles to construct pile bents for a viaduct across coastal marshlands or a shallow bay, where the flexural strength and durability of the concrete cylinder section provides a simple and repetitive means of constructing the bridge without the need for cofferdam structures and footings in the water. A marine environment with water access to the site is con- ducive to delivery and installation of large concrete cylinder piles, and the durability and corrosion resistance of these piles provides advantages in such an environment. Steel LDOEPs have been used nationwide on transporta- tion structures where the relative ease of installation com- bined with the ductility and flexural strength of these piles provide an advantage over alternative foundation types. A typi- cal application for steel LDOEPs may be where an extreme
9 event loading such as a seismic event or vessel collision results in high foundation loadings. Several such examples are described in subsequent sections of this report. Historically, the use of LDOEPs for offshore oil platforms provides a reference base for design of steel pipe piles, particu- larly for long friction piles in clay. Much of the understanding of the behavior of LDOEPs during driving and subsequent axial loading has come from the literature surrounding the off- shore industry (Randolph 2003; Lehane et al. 2005b; Stevens 2010; API 2011). However, the use of LDOEPs for transpor- tation structures differs from offshore applications in several ways. More favorable soil conditions for axial resistance most often exist at bridge sites compared with offshore, as offshore conditions often include deep deposits of soft clay. The rela- tive costs, construction equipment availability, and schedule demands are different, as are the water depths and service life requirements. Offshore piles tend to be primarily shaft resis- tance piles so that base resistance is not as crucial, as often the case for transportation structures. Nevertheless, the offshore experiences are of great value in the design of LDOEPs for transportation structures. TYPES AND CHARACTERISTICS OF PILES Prefabricated tubular steel or prestressed concrete cylinder piles represent virtually all of the LDOEPs used in transporta- tion structures. The manufacture, specification, and handling of these two pile materials are described here. Steel Pipe Piles Tubular steel LDOEPs are formed of steel plate that is bent into a tubular shape and welded to form the pipe. The most economical process for manufacturing typical pipe for use as LDOEPs is the spiralweld pipe, made by using a long coiled sheet that is twisted into a spiral and welded along the spiral seam in a continuous process. The ends of successive coils are straightened and joined before spiraling, and so the spiral- weld pipe comes from the mill as essentially an endless pipe that is cut to individual pipe length (Figure 1). Alternatively, large diameter rolled and welded pipe can be made by rolling plate steel and welding the ends of the plate to form a tube, and then joining a series of individual tubes together to form a long pipe. Spiralweld pipe is typically available in sizes up to 10 ft in diameter, with steel thickness of up to 1 in. and is most often specified by grade with reference to ASTM A252 (ASTM 2010). Grades 1, 2, and 3 have specified yield strength of 30, 35, and 45 ksi, respectively. A252 Grade 3 (modified) can also be obtained with yield strength of 50 to 80 ksi. Thickness of spiralweld pipe larger than 1 in. is not com- mon because of spiral mill capabilities; therefore, rolled and welded pipe is more typically used for greater steel thickness and larger diameters. The seams of spiralweld pipe can be welded from both inside and outside in the manufacturing process and can achieve a full penetration weld that has a strength no less than the steel coil material itself. However, it should be noted that A252 does not specifically require a full penetration weld; therefore, an added note in the agency specification is required to ensure that the spiralweld pipe is manufactured with full penetration welds at the seams. The A252 specification also includes a somewhat generous tolerance on permissible variations in weight and dimensions relative to currently available manufacturing tolerances such that the weight of a pile can be as much as 5% under the speci- fied weight. This difference represents a significant quantity on a large contract. Therefore, a cost conscious manufacturer would read this requirement as a minimum weight that is 95% of the specified weight and supply the material accordingly. As a result, some agencies specify a minimum pile weight rather than defaulting to the A252 standard in this respect. FIGURE 1 Spiralweld pipe: straightening the coil (top); note coils in background; and welding the seam (bottom) (courtesy: Skyline Steel).
10 It is also worth noting that the ends of spiralweld pipe sections represent a cut in the pipe from the process of manu- facturing a continuous pipe. When sections are to be subse- quently spliced together in the field, it is advantageous to mark the cut ends so that these pieces can be re-joined at the same location and thus provide a better fit for field welding. Rolled and welded pipe is less efficient to produce on a large scale than spiralweld and piles produced using this method typ- ically are more expensive on a material basis. This process is often employed for manufacturing pipe for thicknesses that exceed 1 in., or very large diameter pipe as might be used for offshore piling or drilled shaft casing. Steel plate is rolled to produce a tubular shape as shown in Figure 2 and then welded at the ends to produce a straight seam. As the individual cans are welded to form the pipe pile it is a relatively simple matter to modify the wall thickness with length so that greater wall thickness can be provided where needed. For instance, very long piling for offshore platforms often incorporate a greater wall thickness in the upper portion of the pile where flexural strength demand is greater. It is also possible to include a thicker bottom when driving LDOEPs to bear on rock. Corrosion resistance is an issue with steel piles, especially where exposed to air and/or water above the ground surface. The splash zone or tidal fluctuation zone in a salt water envi- ronment is a particularly harsh environment for corrosion of exposed steel piling. Below the soil surface, the presence of high chloride, sulfate ion concentration, or low measured soil resistivity represents aggressive environments for corrosion of steel piling. The 2006 FHWA reference manual on driven piles (Hannigan et al. 2006) provides a summary of current practices with respect to corrosion of steel piling. NCHRP Report 408 summarizes a research study of corrosion of steel piling in non-marine applications (Beavers and Durr 1998). The current standards for the evaluation of this subject are given in AASHTO Standard R 27-01 (2010), which provides a recommended assessment procedure for evaluating corrosion of steel piling in non-marine applications. Corrosion assess- ment for exposed steel piling in marine environments requires evaluation by corrosion specialists. The most common means of addressing corrosion with steel LDOEPs is to provide some allowance of steel loss over the design life of the structure. Coatings can be considered, although the survival of a coating through the pile handling and driving process may pose a challenge. For the portion of steel piling that may be exposed above the soil surface it may be possible to remove soil from within the steel LDOEP and fill the interior with reinforced concrete to a suitable depth below grade. Prestressed Concrete Cylinder Piles The use of prestressed concrete cylinder piles for trans- portation structures has been concentrated along the Gulf and Atlantic coasts, with a typical wall thickness of 6 to 6.5 in., available in sizes ranging from 36 in. to 66 in. out- side diameter. A typical application might be to use these FIGURE 2 Rolling plate for pipe (left) and joining cans (right) (courtesy: Skyline Steel).
11 piles to construct pile bents for a viaduct across coastal marshlands or a shallow bay, where the flexural strength and durability of the concrete cylinder section provides a simple and repetitive means of constructing the bridge without the need for cofferdam structures and footings in the water. A marine environment with water access to the site is conducive to delivery and installation of large cylinder piles. These piles are often manufactured using a spun-cast pro- cess (Figure 3), whereby concrete with zero slump is placed within a rotating drum and spun to produce a very dense and durable concrete with low w/c ratio. Concrete with 8,000 psi compressive strength is routinely produced with this process and the high-strength post-tensioning strands may be used to prestress the concrete to 1500 psi or greater. Confinement of the strands is typically provided by a spiral wire, and stain- less wire may be used in the portion of the pile subject to marine environment. The concrete cylinders are typically 8-, 12-, or 16-ft long sections that contain ducts for post- tensioning after the concrete has cured. The sections are assembled, post-tensioned, the ducts grouted, and the cable ends trimmed at the precast yard prior to shipment. An adhesive is typically used (in combination with the post-tension forces) to seal each joint before pressure grouting the ducts. At the time of this writing (2014), there are two known facilities produc- ing these piles: Gulf Coast Prestress in Pass Christian, Missis- sippi, and Bayshore Concrete Products Corp. in Cape Charles, Virginia. Cylinder piles fabricated using this technique have been installed in one piece with lengths exceeding 200 ft, the primary limitation being the availability of a crane to lift, set, and drive the pile. In recent years, some cylinder piles have been bed-cast in cylindrical forms using an interior form to create the void. Although the concrete does not benefit from the density achieved by the spun-cast technique, the bed-cast method has been used to fabricate piles with wall thickness up to 8 in. and thus achieve greater cover. One challenge with this method of fabrication is to maintain alignment on the interior forms and achieve good filling in the spaces within the formwork. Bed-cast cylinder piles may be prestressed in a manner similar to conventional concrete piling. It may be noted that 36-in.-square prestressed concrete piles are cast in the same way with a 24-in. central void to form a type of LDOEP that has a square outer shape with a cylindrical void. Corrosion resistance of concrete cylinder piles is gener- ally considered to be very good, especially with the spun-cast concrete process, so long as pile damage is avoided during installation. Lau (2005) summarized the examination of three 40-year-old cylinder pile-supported bridges in Florida and found only minor or no corrosion distress of the spiral reinforcement or strand in the piles, in spite of small clear concrete cover values of only 0.4 to 1.5 in. Additional dura- bility is provided by the grouted ducts surrounding the post- tensioning strands and stainless spiral confinement wire can also be used. The avoidance of cracking caused by pile driving appears to be a major factor in achieving durability with concrete cylinder piles. FACTORS AFFECTING DESIGN AND AXIAL RESISTANCE This section provides an overview of those factors affecting the behavior of LDOEPs that are different from conventional smaller piling used in transportation structures. Besides the large diameter compared with most conventional piling, the FIGURE 3 Spun-cast concrete cylinder piles; during casting (top; courtesy: Gulf Coast Prestress), after casting and still in forms (center) and during post-tensioning (bottom).
12 uncertainty in behavior associated with the soil plug within the pile during driving, testing, and subsequent static loading represent challenges that are unique to LDOEPs. Installation and performance of prestressed concrete LDOEPs present some unique conditions relative to steel, and an overview of these features is briefly described at the end of this section. The installation of a large diameter pipe pile engages soil resistance to penetration on both the outside and the inside of the LDOEP. When a large pipe pile is driven into the soil, the hammer imparts a compression wave onto the pipe, which accelerates the pipe downward relative to the soil. For the pile to penetrate during the blow, it must overcome the fric- tional resistance at the pile/soil interface along the outside wall of the pipe. The soil within the inside of the pipe also resists the downward forces exerted by the pipe at the interior pile/soil interface, not only because of the base resistance near the pile toe, but more importantly because of the iner- tial resistance of the soil mass within the pipe. A simplified explanation of this effect is provided here. A Simplified Examination of the Dynamic Behavior of a Soil Plug To understand the behavior of an LDOEP during installation and axial loading it is important to consider the behavior of the soil plug within the interior of the pile. Although there have been numerous papers analyzing the static behavior of the soil plug within a pipe pile, the behavior of the soil plug during installation involves some additional consideration of inertial effects. Because of the inertial resistance of the soil plug to downward acceleration, it is common that an LDOEP may advance without plugging during installation even though the pile may behave like a fully plugged pile during static loading, as illustrated on Figure 4. The pipe is accelerated downward by the action of the hammer. The soil inside the pipe feels side resistance from the pipe as it moves downward and even without any force from below the inertia of the soil mass resists the forces applied by the pipe. For a unit length of pile, the side resistance force, qs is: q d fs i s= pi (1) Where: di = inside diameter of pile, and fs = unit side resistance at pile/soil interface on the inside of the pile. The mass, m, of the soil plug per unit length is: m d g i t = pi γ 4 (2) 2 Where: gt = total unit weight of soil, and g = acceleration of gravity. Assuming zero net force acting on the top and bottom of the plug, the soil plug will then slip when the acceleration of the pile, aslip, is such that: F maslip= (3) And therefore: pi pi γd f d g ai s i t slip=  ï£ï£¬   2 4 4( ) Rearranging Eqs. 2â4 to solve for acceleration, aslip: a g f d slip s i t = γ 4 (5) A typical value for total unit weight is 0.125 ksf, therefore, a g f d slip s i â 32 (6) Where: fs is in ksf and di is in ft. Side resistance outside pipe Inside soil (not plugged) Inside soil (plugged) Side resis- tance inside pipe Base resistance plug + pipe Base resistance pipe only Downward movement pipe only Downward movement plug + pipe FIGURE 4 Schematic of a soil plug inside a pipe pile.
13 This equation provides the simple results illustrated in Figure 5. Since measurements indicate that the acceleration of a large diameter steel pipe pile during driving is likely to be higher than 30 g [Stevens (1988) reported accelerations averaging 178 g] and the unit side resistance on the inside would rarely be expected to be as high as 3 ksf, it is logical to expect that pipe piles larger than 3-ft diameter would rarely be expected to plug during driving. The larger the diameter, the less likely the pile will plug since the acceleration to cause slip is lower with increasing diameter. The obvious conclusion of this simplified analysis is that plugging is unlikely for large diameter driven pipe piles. This conclusion is consistent with a point made in the 2003 Rankine Lecture by Randolph who noted: âthe observations that, under dynamic conditions of pile driving, the soil plug does indeed appear to progress up the pile, with only small variations in the position of the top of the soil plug relative to the original ground surface.â As a point of reference, consider that the area ratio (ratio of the pile cross section to the area within the outside diameter of the pile) of a 48-in.-diameter steel pipe pile with a ¾ in. wall thickness or a 72-in.-diameter steel pipe pile with a 1.125-in. wall thickness is around 6%; this value is less than that of a suitable thin-walled tube sam- pler according to ASTM D1587 (ASTM 2012) (about 8.5% for a 3-in. Shelby tube) used to obtain âundisturbedâ soil sam- ples for laboratory testing. For static loading (zero acceleration), there is no inertial resistance to plugging and the only possible mechanism to push the soil plug into the pipe would be the base resistance mobilized at the pile toe. It can be noted that the simplified analysis of pile plugging provided earlier is only intended to assist the reader in under- standing some fundamental aspects of the problem, and not intended for use in design. The actual dynamic behavior of the plug is more complex than described, because the acceleration and inertia of the pile and plug vary with time as compression and tension waves move up and down the pile (Rausche and Webster 2007). It appears plausible that at least some transient penetration of the plug near the toe can occur during driving if sufficient downward traction is applied by the pile and the base resistance on the bottom of the soil plug is low; for example, a pile penetrating through a sand layer into a clay stratum below might have relatively high internal side resistance as a result of arching near the toe corresponding to low base resistance below the plug. It is clear that plugging behavior in small diameter pipe piles and prestressed concrete LDOEPs (which have thicker walls and smaller inside diameter) may occur under circum- stances in which plugging would not occur for large diameter pipe piles, and therefore observations of pile performance on smaller piles may not properly extrapolate to larger piles. A smaller diameter pile that behaves as a plugged pile during installation may displace a larger volume of soil relative to the pile volume compared with an LDOEP and this condition could influence the unit side resistance, the state of stress in the ground around the pile, the pore pressures generated, and the magnitude and time dependency of setup. Issues Affecting Behavior of Steel LDOEPs During and After Installation A number of other factors have been observed or postulated to have a significant influence on the behavior of steel LDOEPs during and after installation. Many of these are related to the plugging effect and others are related to the size, shape, and length of LDOEPs, as described in the following paragraphs. 0 10 20 30 40 0 0.5 1 1.5 2 2.5 3 ac ce le ra tio n at s lip , a sl ip , g inside unit side resistance, ksf 2ft dia 3ft dia 4ft dia 6ft dia simplified analyses of soil plug behavior, for illustra on FIGURE 5 Acceleration that would cause slip of a soil plug inside an open pipe pile.
14 Base Resistance of Steel LDOEPs on Rock and Driving Shoes The base resistance of driven steel pipe piles has been observed to be relatively low until the pile is installed to bear on rock or a similar hard bearing stratum, suggesting that the base resis- tance is largely dependent on the bearing area of the pile wall itself. Dasenbrock (2006) described observations of unexpect- edly low axial resistance of 42-in.-diameter steel pipe piles that were intended to be driven to bear in sand, with the result that the piles were quite easily driven to refusal to a deeper bedrock stratum to achieve the required nominal resistance. Where steel LDOEPs are driven to bear on rock or other hard materials it is quite common to employ a âdriving shoeâ composed of a ring of steel with greater thickness at the pile toe. If this thickness were to result in a large outside diameter, the effect would be that of a âfriction reducerâ (as might be used above a cone penetration probe to reduce rod friction), with potentially adverse effects on the nominal side resistance of the pile. Most engineers recognize this undesirable conse- quence and therefore use a driving shoe that matches the pile outside diameter and results in a reduced inside diameter. The reduced inside diameter has a similar friction reducing effect on the side resistance within the soil plug and, as a result, the pile is even more likely to drive in the unplugged condition. The long-term impact of this friction-reducing effect may also affect the tendency of the pile to plug during subsequent static loading; however, the use of a driving shoe is generally employed only where the pile is to bear on rock or other hard bearing strata, and so plugging is generally of little conse- quence in such circumstances. The use of a driving shoe that is only a few inches tall may be ineffective in avoiding pile buckling at the toe of steel LDOEPs driven to bear on rock, as evident from Figure 6. A large diameter steel pipe pile driven to bear on rock can quite easily encounter rock on one small portion of the pile toe such that stress concentrations occur on the steel shell. One-dimensional analyses of a pile using wave equation tech- niques to predict pile stresses do not directly account for this non-uniform distribution of stress across the toe. Piles driven through soft soils to bear on a sloping rock surface probably represent the worst possible case for this condition, as a soil plug of weathered rock, till, or even very dense sand may help lessen the risk of buckling to some degree. One effective mitigation strategy that has been employed for steel LDOEPs bearing on rock include the use of a thick- ened bottom section of steel for a length of around 1.5 to 2 pile diameters (M. Holloway, personal communication with D. Brown, Dec. 2013). Another strategy is to âseatâ the pile onto rock using a large number of relatively low energy blows from the hammer in an attempt to achieve more complete con- tact with the rock at the pile toe, followed by only a few hard blows to confirm bearing onto the rock (B. Fellenius, personal communication with D. Brown, Dec. 2013). These strategies have been successfully adopted for installation of 6-ft-diameter steel pipe piles at the new Tappan Zee Bridge after an initial observation of pile damage of a dynamic test pile (Palermo and Reichert 2014). Vibratory Driving and Splicing Where steel pipe piles are used with lengths greater than 100 ft, it is not unusual that a field splice will be required. Splices of steel piling may be accomplished with full penetration welds so that the strength of the splice is equal to that of the pile itself. However, the time required to make the splice may be several hours or more, and so most contractors prefer to stage this work to maintain efficient utilization of pile driving equipment. A common practice is to install the first section of piling with a vibratory hammer and use the impact hammer only to achieve final driving to the required driving resistance. Because the contractor may wish to use the vibratory ham- mer to the maximum extent possible, the agency may be con- fronted with questions related to the hammer requirements for bearing piles and the suitability of the use of vibratory hammers for installation. In general, where steel LDOEPs are installed to achieve the required axial resistance primarily by base resis- tance on rock or a similar suitable hard bearing stratum, the use of vibratory hammers for most of the pile length may not be considered objectionable. Likewise, the uppermost soil strata around a very long pile that may be spliced is likely to be contributing a relatively small proportion of the total side resistance. However, where a substantial portion of the axial resistance is designed to be provided by side resistance in the soil, there is evidence to suggest that vibratory pile installation may result in lower axial resistance (Briaud et al. 1990; Mosher 1990; Canivan and Camp 2002). Most of these comparative studies have been performed on steel H and smaller diameter open-ended steel pipe piles, and at least some of the differences have been attributed to a reduced contribution to axial resis- FIGURE 6 Buckling at the toe of steel LDOEP (courtesy: Bengt Fellenius).
15 tance at or near the pile base. However, it appears plausible that more extensive remolding of soil near the pile wall occurs in cohesive soils with vibratory installation. The vibratory action is thought to liquefy the soil within a narrow zone adjacent to the pile wall during penetration in sands and therefore may result in different relative density and/or horizontal stresses at the pile/soil interface than conventional impact driving. Effect of Pile Length on Behavior and Axial Resistance There are several factors related to the length of pile that may have important effects on behavior. Considerable evidence in recent years (e.g., Randolph 2003; Jardine et al. 2005; Lehane et al. 2005a) suggests that the unit side resistance in both clays and sands can be diminished with increasing length, possibly attributed to: (1) continued shearing of a particular soil hori- zon during pile installation, (2) progressive failure in strain softening soil, (3) reduction in radial stresses with increasing distance above the pile toe, and (4) degradation resulting from densification and/or grain crushing associated with the cyclic shearing action of pile installation. These effects are incorpo- rated to varying degrees in some of the methods for estimat- ing static resistance used for offshore piling summarized by Jeanjean et al. (2010). It is noted that Karlsrud (2012) holds a contradictory opinion (for clay), concluding that pile length or flexibility does not appear to affect the local ultimate shaft friction in clay. A long pile can become quite flexible during compres- sion loading or impact driving, such that the pile undergoes large vertical displacements. The elastic compression of a steel pipe pile that may be 150 to 200 ft in length could easily result in 1 to 2 in. of displacement at the pile top relative to the base, and the long travel time for a compression wave during driving can result in a large number of cyclic stress reversals. At a given point in the soil, perhaps 100 ft below grade, a 180-ft-long pile will result in the soil at the pile/soil interface at that elevation having been subjected to a large number of cyclic stress reversals associated with the penetration of the pile 80 additional feet beyond that elevation. The effects of these many cycles of stress reversal appear to contribute to strain softening behavior at the interface, possibly as a result of degradation of clay soils to residual shear strength condi- tions. If the unit side resistance at the pile/soil interface of a long flexible pile exhibits strain-softening behavior, then progressive failure along the length of the pile can occur dur- ing static loading and the axial resistance degrades toward a residual condition (Randolph 2003). Reduction in radial stresses with increasing distance above the pile toe has been documented in experiments on jacked piles by Lehane and Jardine (1994). Similar behavior in sands is described by White and Lehane (2004), who refer to the effect as âfriction fatigueâ and conclude that the primary mechanism controlling friction fatigue is the cyclic history imparted to the soil elements at the interface during pile installation. Karlsrud (2012) reviewed data from a wide range of pile load tests in clay and concluded that open-ended steel pipe piles generated lower earth pressures against the pile than did closed-ended piles and that reduction of radial effective stress could occur with consolidation of soil around the pile, as excess pore pres- sures dissipate. Experimental evidence of the degradation of sands has been documented by Yang et al. (2010), suggesting that the effects of pile installation on side resistance in sands extend beyond changes in radial stresses and relative density. Calcareous sands can be notoriously brittle and subject to grain crushing, and combined with the low relative density and cementation can result in very low axial resistance of open-ended steel pipe piles as documented by Murff (1987). Time-Dependency of Axial Resistance A time-dependent increase in axial resistance (setup) is known to occur with LDOEPs as is generally the case with all types of driven piles. The increase in axial resistance is affected by soil type, the volume of soil displaced during pile driving, and many other factors. Most of the data available on pile setup are based on smaller piles or displacement piles and, there- fore, some differences in the time dependency of LDOEPs are likely when compared with experiences with other pile types. It is widely accepted that soil disturbance, pore pressures, and time required for consolidation around the pile increase with increasing pile diameter, but that large diameter steel pipe piles may not displace a large volume of soil if plugging does not occur. Relaxation of arching around the pile, cementation, and other ageing effects in the soil at the pile/soil interface extend beyond simple dissipation of pore pressures and con- solidations and result in time-dependent strength gain and setup (Axelsson 2000). Measurements of setup are often based on repeated dy namic tests over time, sometimes even on the same pile. Since LDOEPs tend to have high capacity, it can often be the case that the driving system may be at or near the limits to mobilize the axial resistance and thus the full setup may not be measured (Stevens 2004). Repeated dynamic tests on the same pile can produce degradation issues at the pile/soil interface in calcareous sands, as noted previously, with the result that the measurement of setup is adversely affected by the pile testing history. Driving Resistance and Dynamic Load Testing The use of driving resistance as an indication of pile axial resis- tance has a long history in foundation engineering, and the development of dynamic testing techniques based on stress wave measurements have come into wide acceptance over the last 30 years. However, LDOEPs present some unique
16 challenges compared with the use and interpretation of con- ventional dynamic measurements. The modeling of the behavior and inertial resistance of the soil plug within a large diameter pipe pile during driving pre- sents a challenge unique to LDOEPs. Conventional practice is often based on the simple assumption that base resistance acts only on the annular base of the pile and that the internal and external side resistance are lumped together and consid- ered as external side resistance (Randolph 2003). However, the response of the soil plug is different than that of the exter- nal soil as the inertial mass of the soil plug affects the mobi- lization of internal resistance. Paikowsky and Chernauskas (2008) describe techniques for modeling the soil plug within a pipe pile. The potential for differences in the behavior of the soil plug during impact driving compared with static loading has been described previously, and dynamic testing mea- surements are subject to the same differences compared with static pile behavior if the pile does not plug during a dynamic blow but plugs during static loading. Ordinarily one would expect that the static base resistance of a plugged pile in soil could contribute significant axial resistance that might not be observed during dynamic testing. For long friction piles with low base resistance, it is feasible that the contribu- tion of side resistance derived from the inertial resistance of the soil plug could result in a greater contribution to axial resistance than that of the base resistance from a plugged pile during static loading (M. Holloway, personal communication with D. Brown, Dec. 2013). Static load testing provides the most direct means to measure the static behavior of a test pile, but with the magnitude of loads required to test high- capacity LDOEPs the costs of static tests are very high. The use of a longer duration and lower g force pulse such as the rapid load test method offers advantages in that the pile is more likely to exhibit plugged behavior during a test with lower inertial forces (Muchard 2005). There have been a few attempts to promote plugging within LDOEPs by the use of a partial steel plate within the pile so that the plate allows water through, but engages the soil at some point and presumably makes the piles drive as a full displacement pile. Another factor affecting dynamic measurements on long piles is the potential for residual stresses to affect behavior. Residual stress analysis in the wave equation analysis of piles is fairly well established, but not often performed in routine practice. However, where long, slender, and relatively flexible LDOEPs are analyzed, the use of residual stress analysis is significantly more realistic than a standard model (Rausche et al. 2010). For long LDOEPs, where high base resistance is achieved on rock or other hard bearing strata, it may be quite easy to drive the pile to achieve good bearing on the rock, but the axial resistance that can be observed during dynamic testing can be limited by the ability of the hammer and driving system to mobilize the resistance. This limitation of the hammer can result in a misinterpretation of a dynamic measurement on a restrike blow to conclude that relaxation at the pile toe has occurred when the reality is that the setup in side resistance has diminished the energy reaching the pile toe and thus the mobilized base resistance is reduced upon restrike. In such a case, superposition may be justified (Hussein et al. 2002). Dynamic measurements can be particularly valuable with respect to the detection and avoidance of pile damage during installation. With large diameter piles, hammer alignment on the top of a pile can be more of a challenge. Steel LDOEPs can easily be overstressed at the pile toe when driven to bear on an uneven rock surface as discussed previously. Dynamic measurements can be helpful in detection of damage at the toe of a steel pile; however, damage at the toe is notoriously difficult to detect right away because the reflection from the damage returns at almost the same time as the reflection from the pile toe anyway. Dynamic measurements can identify the onset of a strong base resistance as the pile toe encounters rock, and this identification can be very helpful in controlling the hammer operation to avoid damage. Issues Affecting Prestressed Concrete LDOEPs During and After Installation Although many of the issues described previously apply gen- erally to all LDOEPs, there are some specific issues unique to prestressed concrete LDOEPs, as described in the following paragraphs. Pile Volume and Prestressed Concrete LDOEPs The magnitude of the soil displaced by the pile has an effect on the axial resistance, particularly for piles installed in sandy soil profiles, and even an unplugged prestressed concrete LDOEP may displace a considerable volume of soil. As mentioned pre- viously, the area ratio (ratio of the pile cross section to the area within the outside diameter of the pile) of a 48-in.-diameter steel pipe pile with a ¾-in. wall thickness is around 6%, which is less than that of a suitable thin-walled tube sampler accord- ing to ASTM D1587 (about 8.5% for a 3-in. Shelby tube) used to obtain âundisturbedâ soil samples for laboratory testing. However, a 54-in.-diameter prestressed concrete cylinder pile has an area ratio of almost 40% and displaces 6.3 cubic ft of soil per foot of pile, even if it drives without plugging. This suggests that there could be significant differences in the frictional resistance behavior of these two types of LDOEPs, particularly in sandy soils. There are potentially some significant differences in the behavior of concrete piles compared with steel in the behav- ior of the soil plug during driving. The downward accelera- tion of a concrete pile is generally much lower than that of a steel pile, the volume of soil displaced by the pile wall of a concrete cylinder is much greater, and the diameter of the
17 void smaller. The potential for sandy soil arching within the void is greater (owing to the large area ratio and displaced soil volume), potentially increasing the interior side resis- tance at the pile/soil interface within the plug. McVay et al. (2004) describe analysis of the potential plugging behavior in prestressed concrete cylinder piles similar to the discus- sion of the preceding section; however, these analyses lead to the conclusion that plugging of concrete LDOEPs dur- ing driving is a relatively unlikely occurrence. Rausche and Webster (2007) describe dynamic analyses using wave equa- tion methods including plug soil mass, but conclude that soil plugs often do not develop during driving because of both the plug inertia and lack of internal friction. The large volume displaced by prestressed concrete LDOEPs has been observed to result in the bulking of soil within the pile void such that it was necessary to remove material during installation. Kemp and Muchard (2007) describe longitudinal cracking in Florida believed to be from excessive hoop stress induced by mud and water buildup with the pile during driving, commonly referred to as âwater hammer.â Rausche and Webster (2007) and Muchard et al. (2009) describe occurrences of rising mud and water within prestressed concrete cylinder piles that necessitated removal of the hammer to remove soil from the pile interior; this issue was eventually mitigated in one case by predrilling. Given the large volume of soil displaced by the pile wall of a prestressed concrete cylinder pile and the observations of bulking within the soil plug, it would appear that any cohesive soil within the plug is likely to be very much remolded by the pile driving process. Base Resistance of Concrete LDOEPs Concrete LDOEPs are not commonly driven to bear on hard rock strata, although there have been occasions in which con- crete cylinder piles have been installed onto soft rock or hard bearing layers using a steel driving ring (M. Saunders, personal communication with D. Brown, 2011). This attach- ment was composed of a ¾ in.-thick-steel pipe that extended up through the inside diameter of a spun-cast cylinder pile, protruded 6 in. beyond the end of the pile, and was equipped with a flange to cover the end of the concrete with holes for the post-tensioning strands. Most applications of concrete LDOEPs have been in soil with the pile designed as a long friction pile or else with bear- ing on a dense sand or weak limerock layer. For piles bearing in sand, driving aids such as jetting and/or predrilling are often employed to achieve penetration below the depth required to achieve lateral resistance. Since the area ratio of concrete piles is so large, it is often difficult to distinguish whether a test pile achieved base resistance by behaving as a plugged section or simply through the base resistance mobilized on the pile cross section. In many cases, it may be that the dis- placement required to mobilize the base resistance on the full plugged section may be so large that it is not observed during testing. An interesting comparison was reported by S&ME (2008) between a pair of 54-in. diameter by 80-ft-long concrete LDOEPs bearing in a fairly dense calcareous sand near the South Carolina coast. Both piles were driven open-ended, but one of the piles had the soil plug removed over a large por- tion of its length and replaced with a concrete plug (although the concrete did not extend to the pile toe). The results of the load testing program detected no significant difference in the measured axial resistance between the two piles. Driving Resistance and Dynamic Load Testing Prestressed concrete LDOEPs are typically installed to a specified driving resistance and have many of the same issues related to general installation and dynamic testing as described previously. However, concrete piles have some additional unique considerations during installation in order to avoid potential damage to these piles. Drivability analyses and dynamic measurements are effec- tively used to select pile hammers and cushions for concrete LDOEPs so that high tensile stresses can be avoided (Kemp and Muchard 2007; Rausche and Webster 2007). High tensile stresses can occur with high energy blows when relatively low base resistance is mobilized, particularly with hammers hav- ing a high-impact velocity (such as diesel or some hydraulic hammers). As with any prestressed concrete pile, the driving energy therefore needs to be managed as part of the installa- tion criteria. The hammer must be carefully aligned onto the pile to avoid uneven stresses at the top of the pile that could produce localized overstress or spalling at the pile top, and the use of dynamic measurements with at least four gauges at 90° intervals around the pile can be helpful in verifying good hammer alignment. Hoop stresses in the pile wall may also be present as a result of radial stress from the soil plug (or water hammer, if a sufficient air void at the pile top is not main- tained), and thus the spiral transverse reinforcement in the pile wall is an important component of the pile reinforcing. Prestressed concrete cylinder piles are not typically spliced during installation and then subjected to additional driving. Because driving splices are not commonly used, the maximum length of these piles is limited by the contractorâs ability to lift and drive a pile. For typical transportation structure projects this limits the maximum length to approximately 160 ft. Structural Connection to the Top of an LDOEP In general, the structural connection of the top of the LDOEP to the pier cap or footing is accomplished by installing a reinforcement cage into the pile void and casting a plug of concrete. This approach avoids the large obstruction caused by the extension of the pile wall into a footing or pier cap.
18 The depth of the concrete plug is controlled only by the need to achieve load transfer from the structure to the pile itself, although many agencies prefer to use concrete filling to a depth of a few feet below the scour elevation. This require- ment is likely to necessitate the excavation of the soil plug within the pile to the appropriate depth. For a steel LDOEP, the steel pipe itself may be structurally connected to the footing, as is illustrated by Figure 7. This detail is from the Lafayette Bridge across the Mississippi River in Minneapolis, for which the main piers are each founded on a 4 à 6 group of 42-in.-diameter steel pipe piles. The longitudinal reinforcement in the splice is composed of 20 number 8 bars extending into the pile cap. Figure 8 provides a slightly different detail from the Hast- ings Bridge in Minnesota, an arch bridge with piers similarly founded on 3 à 7 groups of 42-in. steel pipe piles. In this case, shear studs are welded to the steel pipe itself to develop the structural strength of the steel pipe at the connection location. Similar connection details are typical for prestressed con- crete LDOEPs, although the diameter of the interior void is typically smaller relative to the outside diameter. Another consideration for prestressed concrete is that the prestress- ing strands of bed-cast piles also require some development length from the end of the pile for development of the full flexural strength of the pile. DESIGN OF LARGE DIAMETER OPEN-ENDED PILES Design for Axial Loading Whereas the nominal axial resistance of a large diameter open- ended pipe pile is determined in the field based on driving resistance correlated with load test measurements, the static computations of axial resistance serve only as a guide to esti- mate the pile length before driving. Where LDOEPs are driven to bear on rock or other hard bearing strata, the pile length is particularly insensitive to the static computations as the final length will be determined by stratigraphy and the selection of driving equipment rather than static analysis methods. However, static computations of axial resistance are always needed to estimate the lengths of piles expected to terminate in soil in advance of construction. In some cases where driv- ing resistance is not relied upon for determination of axial resistance (notably long friction piles in clay soils), the piles may simply be driven to a predetermined embedded length. In such cases, computed static resistance (perhaps correlated to static load tests) may serve as the basis for final design. Static analysis methods outlined in the current AASHTO design specifications (2013) parallel the methods described in the most recent FHWA manual for driven piles (Hannigan et al. 2006). With the exception of the l method from 1972 (which was developed for offshore piling, but is no longer used offshore), none of the empirical methods described in these publications were developed specifically for LDOEPs or even based on data from load-tested LDOEPs. Most recent literature references do not give serious attention to the FHWA and AASHTO procedures for computing nominal axial resis- tance of LDOEPs. It appears that the resistance factors for piles included in the current AASHTO guidelines do not specifically represent LDOEPs. The resistance factors are based largely on the work reported by Paikowsky (2004) in NCHRP Report 507, and the database of load tests used to develop the recommenda- tions for LRFD resistance factors includes a very small num- ber of open-ended pipe piles. LDOEPs are not documented separately from smaller open-ended pipe piles, but logically represent an even smaller portion of the data. The current AASHTO code (2013) does not distinguish the design of deep foundations on the basis of any of the unique characteristics of LDOEPs. The most widely referenced procedure for the design of large diameter open-ended steel pipe piles is the API RP 2GEO (2011) procedures for offshore pile foundations. The American Petroleum Institute (API) also references other possible methods in the commentary. The following paragraphs provide a brief summary of the computational methods described in the literature that are par- ticularly relevant to LDOEPs in soil. Axial Resistance in Clay Soils The side resistance in clay soils determined using the API methods are based on correlations with undrained shear strength, su, using a dimensionless empirical correlation factor, alpha (a), which typically is taken to be 1.0 or less. The API methods differ from other alpha methods in the approach used to determine alpha and the means of estimating su. Undrained shear strength is not an intrinsic material property, but rather a function of the test method used to measure it; in addition, the measurement of su is subject to the effects of sampling disturbance and other factors. Because of the challenges of sampling and testing in an offshore environment, the API pro- cedure includes suggestions for estimating su as a function of Over Consolidations Rules and effective vertical stress (p0â²). This method is presented as follows: f (z), the unit shaft friction can be calculated by: ( ) = α (7)f z su Where: a is the dimensionless shaft friction factor for clays; and su is the undrained shear strength of the soil at the point in question, in stress units.
19 FIGURE 7 Pile to footing connection, Lafayette Bridge, Minnesota.
20 The factor a can be computed by: α = Ï Ï â¤â0.5 for 1.0 (8)0.5 α = Ï Ï >â0.5 for 1.0 (9)0.25 with the constraint that a ⤠1.0, and where: ( )( )Ï = â² at depth, (10)0s p z zu ( )â² = effective stress at depth (11)0p z z Where the pile toe is in cohesive soils, the unit base resis- tance, q, is estimated as equal to 9su, a value that typically represents a low proportion of resistance compared with the side resistance. The side resistance is assumed to act on both the inside and outside of the pile, with the limitation that the resistance on the inside of the pile is limited by the base resistance of the plugged section below the toe. Because the base resistance of piles in clay is relatively low, the interior side resistance does not normally contribute. The API procedure provides discus- sion of the possible reduction in the computed nominal side resistance as a result of pile length effects, low Plasticity Index (PI) clays, and highly overconsolidated soils (y > 3), with reference to the commentary to the API code; however, discretion on these issues is left to the designer. The âalpha methodâ approach to correlating unit side resis- tance with undrained shear strength of clay soil serves as a basis for some additional methods that follow this general methodology. Saye et al. (2013) provides a review of the issues of sample disturbance and the use of the normalized stress history (the âSHANSEPâ approach) to address sam- FIGURE 8 Pile to footing connection, Hastings Bridge, Minnesota.
21 ple disturbance problems affecting side resistance using alpha methods. This method offers a potential way of addressing one of the most common problems with the use of undrained shear strength as a basis for design, namely the contamination of the design soil strength profile with data from âundisturbedâ strength measurements that might be affected by sample dis- turbance. Karlsrud (2012) provided a recent review of avail- able pile static load test data in cohesive soils leading to a modified procedure for estimating alpha as a function of PI to account for some unusually low axial capacity load test data in low PI soils. Axial Resistance in Sands The API methods in siliceous sands are based on the use of a shaft friction factor âbetaâ (b) and end-bearing factor Nq, which are multiplied by the effective vertical stress to obtain unit values of side and base resistance, respectively. This approach is fundamentally the same as the beta method described in the FHWA manual (Hannigan et al. 2006), but relies on a table of specific design parameters that are recom- mended for pipe piles based on the estimated relative density and grain size description of the soil. In general, it is con- sidered that there is higher variability in computed nominal resistance in sands than in cohesive soils; however, dynamic measurements of driven piles in sands is likely to have some- what greater reliability as an indicator of axial resistance; there- fore, the overall reliability of piles in sands is not necessarily lower than for cohesive soils. Methods Utilizing CPT Data Other methods rely on cone penetration test (CPT) measure- ments in soils with adjustments to account for pile length and other effects; a summary of these methods is described in API RP 2GEO (2011). These include the ICP-05 methods pro- moted by the Institute of Civil Engineers (English) (Jardine et al. 2005), the UWA-05 methods promoted by the Australians (Lehane et al. 2005b), the NGI05 methods promoted by the Norwegians (Clausen et al. 2005), and the Fugro05 methods promoted by the Dutch (Kolk et al. 2005). Although a con- sensus approach has not emerged, these methods have many similarities. There appears to be merit and increased interest in the use of these approaches since they generally account for effects not included in the simplified API procedure; most of these methods also rely on CPT data, which may have advan- tages in terms of reliability and stratigraphic coverage relative to conventional methods based on laboratory tests. Methods Specific to Prestressed Concrete LDOEPs Similar methods to those described earlier may be employed for prestressed concrete cylinder piles, although none of these were developed specifically for prestressed concrete piles. McVay (2004) performed a study of axial resistance of cylin- der piles for the Florida DOT (FDOT) that included load tests on 22 prestressed concrete cylinder piles from five separate sites; 19 of the 22 piles were from a site in the Florida pan- handle and two sites in Virginia. McVay developed empiri- cal correlations specifically for prestressed concrete LDOEPs with standard penetration test measurements (N, uncorrected for overburden pressure, and presumably N60, although not stated) based on interpreted unit side ( fs) and base (qt) resis- tance in units of tsf, as follows: f C Ln N Cs ( )= â (12)1 2 q C Nt = (13)3 Where: C1,2,3 = empirical constants for the range 5 < N < 60, as listed in Table 1. Design for Lateral Loading The design of LDOEPs for lateral loading is fundamentally no different than that of any other deep foundation element. Some issues with respect to the pile itself include the structural con- nection to the pile cap or footing described previously and the effect on the foundation stiffness contributed by the potential concrete plug within the upper portions of the pile. Where the pile is filled with concrete, the concrete adds considerable stiffness and likely generates behavior as a com- posite concrete/steel shell member in the same way that a permanent steel casing contributes strength and stiffness to a drilled shaft. In California, concrete-filled steel LDOEPs are sometimes referred to as âCISSâ or cast-in-steel-shell piles. Soil C1 C2 C3 Plastic Clays 0.5083 0.634 0.2226 Clay-Silt-Sand Mixtures 0.3265 0.5404 0.4101 Clean Sands 0.0188 0.0296 0.5676 TABLE 1 EMPIRICAL CONSTANTS C1,2,3
22 There exists some uncertainty about the distance from the ends of the pile required for the interior concrete to develop enough bond for full composite action, and it is understood that there are several research initiatives related to this issue for drilled shaft foundations that may have relevance to the structural behavior of concrete-filled LDOEPs. Design for Settlement/Uplift/Serviceability The design of groups of LDOEPs for settlement is funda- mentally no different than that of any other deep foundation system, and general guidelines for estimating settlement of pile groups are provided by FHWA (Hannigan et al. 2006). The same is true for LDOEPs designed to resist uplift forces. The axial stiffness of an LDOEP can be affected by the rela- tive contribution of the base resistance if plugging behavior is anticipated and significant base resistance of the full plugged pile cross section is considered. Where a large diameter base contributes significantly to the axial resistance, the displace- ment required to fully mobilize that base resistance may be significant. For this reason, some agencies (e.g., Florida DOT) use a modified form of the Davisson offset method for inter- pretation of static load tests on piles that are 24 in. or larger, whereby the displacement at the strength limit is based on an offset of D/30 rather than D/120 as used for smaller piles. This greater value reflects the larger displacement to fully mobi- lize the base resistance. Where a computer model is used to replicate the stiffness of a group of piles, the ât-zâ springs for the base resistance may need to be adjusted based on whether the LDOEP is to reflect the plugged behavior with associated larger displacement needed to fully mobilize the base resis- tance on the plug or whether the unplugged base resistance is anticipated at small displacements. SUMMARY This chapter provides a summary of background information on the use of LDOEPs for transportation projects, outlining some of the important issues to be addressed in the selection and design of LDOEPs for this purpose. LDOEPs may con- sist of steel pipe or prestressed concrete cylinder piles and are defined for the purposes of this report as driven, open- ended piles that are of 36 in. outside diameter or larger. LDOEPs provide advantages where large foundation loads may exist and/or the piles are subject to significant unsupported length as a result of scour, liquefaction, or very weak surficial soils. Marine construction conditions also favor the use of these piles, particularly where pile bents might be employed to elimi- nate footings. Steel pipe is often specified by grade with reference to ASTM A252 (ASTM 2010) and may be economically manu- factured as spiralwelded pipe using a long coiled sheet that is twisted into a spiral and welded along the spiral seam in a con- tinuous process. Where piles larger than 10 ft in diameter or thicker than a 1-in. wall are required, rolled and welded straight seam pipe may be used. Prestressed concrete cylinder piles have advantages of corrosion resistance and durability, which may be particu- larly important for coastal structures. These piles may be fabricated as spun-cast cylinders using low-slump concrete, which results in concrete with high strength (typically com- pressive strengths of 8,000 psi or greater), low permeabil- ity, and high density. The cylindrical sections are assembled and post-tensioned to fabricate piles with a length of up to 200 ft or greater. Concrete cylinder piles have also been cast in conventional horizontal prestressing beds using a form insert to create the center void. A range of factors that are distinctive to LDOEPs as opposed to conventional piles are described, most notably the behavior during driving and the tendency of the interior soil to remain in place or even rise within the pile as the pile is driven. The failure of most LDOEPs to âplugâ during initial installa- tion is related to the inertial resistance of this soil mass as the pile is accelerated downward by the hammer. Plugging may be more likely during subsequent static loading where inertial forces do not contribute, and this difference between behavior during installation and subsequently is the source of some dif- ficulty with the use of driving resistance or even high-strain dynamic load tests as an indicator of static axial resistance. Steel LDOEPs have advantages when driven to bear on rock, because the âunpluggedâ behavior during installation allows the pile to penetrate to the rock with relatively less driving resistance until bearing is achieved. However, high and potentially nonuniform end-bearing stresses at the pile toe require consideration. Installation of steel LDOEPs using a vibratory hammer can provide another advantage for ease of installation, par- ticularly where a splice is required and the vibratory ham- mer may be used to install the first pile section. However, the uncertainty related to the effect of vibratory installation on subsequent axial resistance dictates that steel LDOEPs are typically driven to bear using conventional impact hammers. The time dependency of axial resistance related to setup is a consideration with pile installation criteria as it is with any driven pile; however, LDOEPs have additional uncertainty related to the potential difference in behavior of the soil plug during dynamic penetration and static loading. Issues of dif- ferences between how dynamic and static loading is applied to the pile need to be included when evaluating setup with restrikes and/or load tests.
23 Prestressed concrete LDOEPs have many similar issues related to the interpretation of driving resistance, soil plug- ging, setup, etc.; however, a distinctive feature of these piles relative to steel is the large area ratio of the pile cross section that certainly affects the behavior of the soil plug and the avail- able cross-sectional area of the pile to engage base resistance during installation. Prestressed concrete LDOEPs also require consideration of potential tensile stress during driving, as with any prestressed concrete pile. The design of LDOEPs is distinctive from other types of driven piles primarily in terms of the computation of axial resistance. Although the AASHTO design codes do not dis- tinguish these piles from other types of driven piles, the pile load test data that were used to establish resistance factors for design included very few examples of LDOEPs. The most commonly used computational procedures for estimating static axial resistance of steel LDOEPs in soil are found in the API guidelines, which have a history of use for the design of offshore platforms. The API procedures have expanded in recent years to include additional calculation methods based on the use of CPT and to account for length effects and other factors such as partial plugging. FDOT has spon- sored one study to develop empirical design procedures for LDOEPs based on standard penetration test (SPT) measure- ments, which include data from a few prestressed concrete LDOEPs. In conclusion, this summary of background information has identified many of the distinctive features affecting the design, construction, and testing of LDOEPs, and this infor- mation serves as a base of reference for the subsequent con- sideration of the synthesis of practice from transportation agencies described in the following chapters.
