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Cypress Street Viaduct 70 TRANSPORTATION RESEARCH RECORD 1411' Seismic Pile Foundation Stiffnesses at Cypress Street Viaduct G. NORRIS, R. SIDDHARTHAN, z. ZAFIR, AND P. GOWDA Given the soil borings, the estimated soil properties, the free­ tops and from that the maximum moments in the piles) cannot field ground surface motions (at strong-motion stations on either be made. This is clearly demonstrated in a parallel analysis side of the Cypress Street viaduct in Oakland, California), and of the nearby Oakland Outer Harbor Wharf (2). the assessed buildup in porewater pressure in the Merritt sand discussed in the preceding paper in this Record, the nonlinear variation in the lateral and vertical-rotational pile foundation stiff­ nesses at Bents 61 (in Merritt sand) and 97 (in Bay mud) is LESSONS LEARNED FROM ASSESSED VERSUS established on the basis of the methodology outlined in 1992 by OBSERVED RESPONSE OF OAKLAND OUTER Norris. Assessed lateral behavior is compared with observed re­ sponse from the California Department of Transportation (Cal­ HARBOR WHARF trans) lateral pile group load tests at these bents. Differences between stiffnesses corresponding to conditions reflecting the The California Strong Motion Instrumentation Program Caltrans tests and those applicable during the Loma Prieta earth­ (CSMIP) records for the instrumented Oakland Outer Harbor quake are discussed. Likewise, differences between the meth­ Wharf, located 1.6 km (1 mi) west of the Cypress Street odology used here and that reflected by FHWA and the Applied viaduct, provide very important data for the purpose of under­ Technology Council are pointed out. · standing pile foundation response during an earthquake. Although the piles at the Oakland Outer Harbor Wharf are Structural dynamic modeling of a highway bridge requires different [0.46 m (18 in.) square free-standing precast concrete knowledge of the lateral and vertical-rotational stiffnesses of piles at large spacing] from those at the Cypress Street viaduct, its foundations, which are typically pile foundations. Uncou­ they derive their lateral resistance from a surface layer of Bay pled lateral and vertical-rotational pile group stiffnesses are mud (as do piles at Bent 97 of the Cypress Street viaduct) in turn a function, respectively, of the lateral and vertical and their vertical resistance from point bearing in an under­ responses of the individual piles, which can be evaluated via lying dense sand (similar to the piles at Bent 61 in Merritt so-called p-y and t-z analyses. [See the paper by Norris (1) sand). Therefore, there is something to be gained in the pres­ for a general overview of such analysis, with reference to case ent consideration of foundation response at the Cypress Street studies, for both nonliquefying and liquefying soil conditions.] viaduct from this more detailed study of the Oakland Outer It is such lateral and vertical-rotational pile foundation stiff­ Harbor Wharf. nesses in Merritt sand (at Bent 61) and in Bay mud (at Bent In work undertaken by the authors relative to the Oakland 97) at the Cypress Street viaduct that are sought here. Outer Harbor Wharf (2), it was found that using free-field Unfortunately for the bridge engineer, these stiffnesses are motion input and (assessed) compatible lateral and vertical nonlinear-that is, they are displacement or rotation depen­ pile stiffnesses, a linear structural dynamic model is capable dent-and a linear dynamic analysis of a bridge will require of providing computed motions on the deck of the wharf that the choice of boundary element spring stiffnesses that are nearly perfectly match the recorded accelerations on the struc­ compatible with the resulting relative displacements or ro­ ture. On the other hand, using stiffnesses that are too high tations. This means that given the nonlinear variation in the (e.g., assuming fixed foundations) causes an assessed peak stiffnesses (e.g., at Bents 61and97), the bridge engineer must acceleration (0.58 g) twice the recorded acceleration and shear assume a set of stiffness values for a linear model and repeat and tensile forces at the head of the critical pile almost three the dynamic analysis, supplying a new set of stiffness values times that of the compatible stiffness solution. Similar com­ each time, until the resulting relative displacements yield stiff­ binations of lateral and vertical stiffnesses that are too high ness values that match the assumed set of values. It is only or low (as compared with displacement-compatible values) in this fashion that one would be able to assess the progressive cause poor predicted responses. failure of the Cypress Street viaduct, which occurred over just An important point in the evaluation of the compatible the northern portion of its length. stiffnesses is that the soil modulus values (Young's E or shear Although such analysis will require several iterations, un­ modulus G) employed in such assessment should be chosen less it is undertaken, correct determination of the distribution in relation to the governing strain condition in the soil. At of (seismic) load to· the structure, the acceleration on the present, FHWA (3) and Applied Technology Council (ATC) structure, or the forces in the piles (i.e., the shears at the pile ( 4) recommendations in regard to this matter are at odds. FHWA would have the engineer assess stiffnesses on the basis of traditional p-y and t-z approaches, where pile responses G. Norris, R. Siddharthan, and Z. Zafir, Department of Civil Engineering/258, University of Nevada, Reno, Nev. 89557. P. Gowda, are a function of the relative displacements, as if the soil 25 Poncetta Drive #118, Daly City, Calif. 94015. immediately surrounding the foundation (the near-field soil), Norris et al. 71 -Z-.:supported mass of superstructure near field re<]ion of inertial far field region interactioo, sane as re<]ion of c_;_•• field mponsel lateral pile response H M~ upwardly propagating shear wave --- a) overlapping wedges of developing passive pressure r~sistance cylindrical zones of vertical resistance b) FIGURE 1 (a) Near- versus far- (i.e., free-) field soil regions and (b) differing near-field (or inertial interaction) soil regions for lateral and vertical pile response. as shown in Figure la, were moving and the free-field soil showing a cutoff at 350 kN/mm (2,000 kips/in.) at a relative (also called the far-field soil) were not. During seismic exci­ vertical pile head movement of 0.75 mm (0.03 in.), corre­ tation, this would not be the case. ATC, on the other hand, sponding to the equivalent free-field shear strain of 3 x 10-2 would have the engineer use soil modulus values based on percent in the dense sand and 2 x 10- 1 percent in the over­ the level of the free-field strain, as if there were no relative lying Bay mud. The computer program SHAKE (5) was used motion in the near-field soil (i.e., the near-field soil moves to obtain these shear strains given the free-field ground sur­ in an identical manner to the free-field soil). face acceleration record shown in Figure 11 of the preceding Of course, the appropriate procedure would be to take the paper in this Record. The shear moduli, G, of the soils in the· modulus values in the associated near-field soil region, as t-z. program were then limited to the corresponding values shown in Figure lb, as a function of the total strain (free­ from the SHAKE program, resulting in a constant vertical field plus relative, with due regard to phase differences). pile head stiffness (350 kN/mm) for relative pile head dis­ However, an expedient that would be more reasonable than placements less than 0. 75 mm (i.e., the free-field strain dom­ either the FHWA or A TC approaches would be to use the inates) and a diminishing stiffness at successively greater dis­ larger of the two strains (free-field or relative) in the near­ placements (i.e., the near-field or inertial interaction ·strain field soil region, that is, assume that either one or the other dominates). By contrast, the FHWA approach would have dominates. However, since the governing relative strain is a the stiffness increase at lower displacements ( <0.75 mm), function of the level of the inertial interaction response of the whereas the A TC approach would have the stiffness remain superstructure, the engineer does not automatically know which constant at 350 kN/mm (i.e., a horizontal line) across all levels is the larger. Nevertheless, this approach can be accommo­ of relative displacement. It should be pointed out that in the dated by showing the stiffness, determined as a function of case of the Oakland Outer Harbor Wharf, the compatible the relative displacement or strain, with a superposed cutoff solution resulted in the vertical pile stiffnesses falling (as dif­ where the relative displacement or strain would be less than ferent points for differently loaded piles) on this horizontal the free-field value. cutoff of 350 kN/mm, whereas the compatible horizontal stiff­ Figure 2 is such a plot of the nonlinear variation in the nesses for all piles fell on their appropriate curve, to the right vertical pile stiffness at the Oakland Outer Harbor Wharf of the free-field cutoff. In other words, it was the free-field 72 TRANSPORTATION RESEARCH RECORD 1411 Deflection (in) 0 0.1 0.2 0.3 0.4 0.5 E' 600 CD< E ;+ .......... 3000 z c:;· 6 Q_ (/) O') r+ O') 400 Cl) ~ c ::J 2000 CD ~ rJ) iii rJ) ........... 0 CJ -o·"'" 200 rJ) t 1000 .......... ~ 2: 0 0 0.0 2.5 5.0 7.5 10.0 12.5 Deflection (mm) FIGURE 2 Vertical pile stiffness variation for 0.46-m (18-in.) square prestressed concrete piles at Oakland Outer Harbor Wharf during the Loma Prieta earthquake.
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