Many different materials and shapes have been used for pile designs. Although numerous examples of long span bridges in seismic areas founded on timber piles still exist in California (Priestley and Seibel, 1997), one being the San Francisco Oakland Bay bridge, the current trends involve the use of concrete (reinforced or hollow prestressed) piles, steel (H type, shell and concrete filled shell) piles; a special situation is represented by the integral pile-shaft column arrangement where the piles are not connected to a pile cap but are extended by columns into the superstructure.
In most countries for which information was collected (Table 5-3), reinforced concrete piles largely prevail (>70% and most of the time >90%). Exceptions are represented by Japan with 50% of concrete piles, half of it being prestressed, and USA (California) with only 30% of concrete piles. When steel piles are used they are either concrete filled shell or shell piles with an exception for France where 13% of the steel piles are of the H type. Integral pile shaft columns seem to be used almost exclusively in California and to a much lesser extent in France (respectively 60% and 5% of the total number). Although not explicitly quoted in Table 5-3, prestressed solid piles are sometimes used in California with a square or octagonal cross section, with diameter between 400 and 600 mm.
With regards to the installation techniques almost all the concrete piles are cast in drilled holes, except in Japan and California where 30% of the concrete piles are driven. Steel piles are mainly cast in place (Table 5-3).
USA
-
California France Gree
ce
Italy Japan New
Zealand Slovenia Taïwan
Integral pile shaft columns 60 4 H type 4 15 Shell 2 5 50 (1) Steel piles Concrete filled shell 10 20 50 (1) 15 Reinforced 30 70 ~100 95 50 (2) 70 100 99 Pile type Concrete piles Prestressed shell 50 (2) 1 Cast in drilled holes 70 69 ~100 X 70 60 100 99 Concrete piles Driven 30 6 30 20 1 Cast in shell piles 20 X 70
Installation procedure Steel
piles
Driven 5 30 20
(1) percentage with respect to steel piles (2) percentage with respect to concrete piles
One controversial topic regarding the pile type relates to the possibility of using raked piles in a seismic environment. From a static point of view, this type of foundation may be advantageous for abutments to sustain the horizontal forces induced by the static earth pressures. However for seismic loading, several codes, like for instance Eurocode8, advise against the use of raked piles. This is based on the observation that these piles frequently suffer substantial head damage during earthquakes. This occurs when the axial stiffness of the pile is much greater than the lateral stiffness (Pender, 1993). In this situation even a small angle of rake gives horizontal and moment stiffness terms of the pile head stiffness matrix considerably larger than those for the vertical pile. The consequence is that when a horizontal displacement load is applied at its head the pile sustains large axial forces; the behavior under this type of loading is obviously less than ductile than when the pile responds mainly in flexure. In addition, during earthquake loading the supporting soil may undergo vertical settlements causing a lack of underneath support to the inclined piles, which still carry the overburden load; these parasitic bending moments must be added to the axial forces. All these issues are design issues that may be accommodated provided the physics of the phenomena are perfectly understood, but this has not always been the case. There is however one situation for which inclined piles seem to have a beneficial effect: when there is potential for liquefaction and lateral spreading at a bridge site the use of inclined piles help reducing the damages to the foundation by limiting the horizontal displacements as shown by the example of the Edgecumbe bridge (Berrill et al, 2000). Additional means for enhancing the resistance to lateral displacements of the foundation system have been presented in section 3.2.
5.3.2 Modeling techniques
As opposed to spread footings, for which a single method of analysis to determine the forces transmitted by the foundation emerges in practice (based on a substructuring approach and the definition of the foundation stiffness matrix and damping), several modeling techniques are used to model pile foundations in a global bridge model for seismic response studies; the most common methods are the simplified beam on Winkler foundation model and the coupled foundation stiffness matrix (substructuring). These two modeling techniques are illustrated in Fig.5-3 for the complete model and in figure Fig.5-5 for the substructure model (Lam et al, 2004).
In the complete model, piles are represented by beam elements supported by linear or non linear, depth-varying, Winkler springs. In the case of earthquake excitation, ground motion would impart different loading at each soil spring and these motions need to be calculated from a separate analysis.
The main drawback of this modeling technique is the large number of degrees of freedom needed to formulate the complete system. The alternative approach employing a substructure system in which the foundation element is modeled by a condensed foundation stiffness matrix and mass matrix along with equivalent forcing function represented by the kinematic motion, may be more attractive; in addition, it more clearly separates the role of the geotechnical engineer and of the structural engineer. The substructuring approach is based on a linear superposition principle and therefore linear soil behavior is more appropriate. In that case, the condensed stiffness matrix may be obtained either from the beam on Winkler springs model or from continuum impedance solutions (Gazetas, 1991). When non linear soil behavior is considered, the condensed stiffness matrix is generally evaluated by a pushover analysis of the pile group and linearization at the anticipated displacement amplitude of the pile head.
Fig.5-3: Complete pile-structure model
The p-y relation, representing the non linear spring stiffness, is generally developed on the basis of a semi-empirical curve, which reflects the nonlinear resistance of the local soil surrounding the pile at specified depth. A number of p-y models have been proposed by different authors for different soil conditions. The two most commonly used p-y models are those proposed by Matlock (1970) for soft clay and by Reese et al (1974) for sand. These models are essentially semi-empirical and have been developed on a basis of a limited number of full-scale lateral load tests on piles of smaller diameters ranging from 0.30 to 0.40 m. To extrapolate the p-y criteria to conditions that are different from the one from which the p-y models were developed requires some judgment and consideration. For instance in Slovenia, values of the spring stiffnesses are derived from the static values, increased by 30%. Based on some field test results, there are indications that stiffness and ultimate lateral load carrying capacity of a large diameter drilled shaft are larger than the values estimated using the conventional p-y criteria as reported by Stevens and Audibert (1979). Pender (1993) suggests that the subgrade modulus used in p-y formulation would increase linearly with pile diameter. Using nonlinear three dimensional finite element analyses, Lam and Law (1996) demonstrated that the increases in stiffness and lateral capacity of large diameter shafts are attributed to additional soil resistance mobilized due to pile rotation.
Studies have shown that Matlock and Reese p-y criteria give reasonable pile design solutions. However, the p-y criteria were originally conceived for design against storm wave loading conditions based on observation of monotonic static and cyclic pile load test data. Therefore, Matlock and Reese’s static p-y curves can serve to represent the initial monotonic loading path for typical small diameter driven isolated piles. If a complete total system of a bridge is modeled for seismic response study, individual piles and p-y curves can be included in the analytical model. However, for a large pile group, group effects become important. An example is given in Fig. 5-4 which presents the results of horizontal impedance calculations of the group of piles of half the foundation (22 piles) of one of the pylon of the Vasco da Gama bridge in Lisbon (Pecker, 2003); the group efficiency, computed from elastodynamic theory, is of the order of 1/6 at low frequencies and decreases with frequency due to the constructive interference of diffracted waves from adjacent piles. Typically, for large pile groups it is not uncommon to calculate group efficiency in the range 1/3 to 1/6.
5.80 m 5.00 m 5.80 m 20.20 m 41.95 m 44 piles φ = 2.20m 5.80 m 5.00 m 5.80 m 20.20 m 41.95 m 44 piles φ = 2.20m 0 250 500 750 1000 0.0 0.2 0.4 0.6 0.8 1 Frequency (Hz) Re a l p a rt (M N/ m ) Isolated pile Pile group .0
Fig. 5-4 : Horizontal pile group impedance for the Vasco da Gama bridge (Pecker, 2003)
Although the group effects have been a popular research topic within the geotechnical community, currently there is no common consensus on the design approach to incorporate group effects. Full scale and model tests by a number of authors show that in general, the lateral capacity of a piles in a pile group is less than that of a single isolated pile due to so- called group efficiency. The reduction is more pronounced as the pile spacing is reduced. Other important factors that affect the efficiency and lateral stiffness of the pile are the type and strength of soil, number of piles, type and level of loading. In the past, the analysis of group effects were based mostly on elastic halfspace theory due to the absence of costly full- scale pile experiments. In recent years, a number of major studies yielded some high quality experimental data from full-scale or centrifuge model tests (e.g., Ashford et al, 1999, Brown, et al, 1987, McVay, et al, 1995). In addition to group effect, gapping and potential cyclic degradation have been considered in the recent studies. It has been shown that a concept based on p-multiplier applied on the standard static loading p-y curves works reasonably well to account for pile group and cyclic degradation effects (e.g., Bogard and Matlock, 1983; Brown et al, 1987; Brown and Bollman, 1996). The p-multiplier is a reduction factor that is applied to the p-term in the p-y curve for a single pile to simulate the behavior of piles in the group. For instance, the values proposed by Brown and Bollman (1996) are given in Table 5-4. Clearly, p-multipliers are dependent on site conditions, soil types, details of stratification and displacement amplitudes (Finn, 2005).
Row spacing
Front row Second row Third and more rows
3D 0.80 0.45 0.35
4D 0.90 0.65 0.55
5D 1.00 0.85 0.75
Table 5-4 : p-multipliers for pile group design (Brown and Bollman, 1996)
Fig.5-5: Substructure model
Examination of the questionnaire filled by the various countries mentioned above show that in practice only one method of analysis is considered for the analysis of pile foundation: the p-y curve method with the exception of France considering both methods (p-y curves and continuum analysis) as viable alternatives (Table 5-5). In all countries, but France, group effect is taken into account. Kinematic interaction, which serves to define the effective input motion is only considered in Japan and, for the verification of the pile integrity, in Greece under certain restrictive conditions (high seismicity, bridge of high importance and soft heterogeneous soil profile).
All countries have a specific design code for bridges, based on Eurocode Part 2 in Europe, and on Caltrans specifications in California or AASHTO outside California in the United States , JRA in Japan, AASHTO and own bridge code in Taiwan.
5.3.3 Pile integrity checks
Because of difficulties in investigating pile conditions after an earthquake, it will be normal to design piles to remain essentially elastic under design level seismic response. An exception to this rule applies for the integral pile shaft-column designs, where development of an in-ground plastic hinge is inevitable when the pile and the column have the same diameter and reinforcement; when the pile diameter is larger than the column diameter, as sometimes encountered in New Zealand, the plastic hinge can be placed at the ground level (Fig. 5-6). However, some codes, like Eurocode 8 (part 5), allow for the formation of a plastic hinge in the pile: "Piles should in principle be designed to remain elastic, but may under certain conditions be allowed to develop a plastic hinge at their heads". Nevertheless, in all countries having filled the questionnaire but one, no advantage is taken from this allowance and the piles are designed to remain elastic and the connection pile to pile cap to be fixed.
Fig. 5-6 : Example of integral pile shaft column used in New Zealand
USA
-
California France Gree
ce
Italy Japan New
Zealand Slovenia Taïwan
Pinned or Plastic hinge 50 30 Column / base Fixed 50 70 X X X X X X Pinned or Plastic hinge 2 Connection Pile / pile cap Fixed X 98 X X X X X X p-y curves X X X X X X X X Continuum Footings
Group effect Yes No Yes Yes Yes Yes Yes Yes
Analysis
Kinematic interaction Yes Yes Yes (1) No Yes No No No
(1) under conditions set forth in Eurocode 8 (EN 1998-5).
Table 5-5 : Methods of analysis for pile foundations
Greece is the only exception to the above rule and allows for the formation of a plastic hinge when elastic design is not possible; under that circumstance, confinement of the concrete core at the potential and possible regions of plastic hinges must be effected, as well as capacity shear check of the piles. Potential region for a plastic hinge is considered a region of length 2d below the pile cap. In addition, if the pile crosses the interface of successive soil
layers which have much different shear moduli (ratio of shear moduli > 5), then regions of ±2d above and below the possible limits of this interface shall be deemed to be regions of possible plastic hinges. In these regions, confinement and bending strength equal to that of the pile top shall be provided. This rule does not apply for the region of the foundation layer for end-bearing piles provided that conditions of full fixity of the piles are not developed there.
For a plastic hinge forming at the pile top with a monolithic pile/cap connection, the plastic hinge length may be taken equal to (Priestley et al, 1996)
0 08c 0 022 y bl 0 044
L= . l + . f d ≥ . f dy bl (5-5)
where lc is the distance from the point of contraflexure to the pile cap, fy is the yield strength
of pile longitudinal reinforcement in MPa and dbl is the longitudinal bar diameter.
For in-ground hinge, the hinge length decreases with increasing soil stiffness and with reducing height from ground to the point of contraflexure. The lower limit is approximately given by :
1 0 0 06
L= . D+ . H D (5-6)
where H is the above grade height to the point of contraflexure. Recent experiments (Budek et al, 1997) have shown that the amount of confinement reinforcement can be reduced in the plastic hinge region because of the beneficial effect of the soil pressure exerted on the compression side of the pile (Fig.5-7); changes in the pile moment profile are caused by solid pressure against the compression side of the pile and results in additional confinement stress being placed on the compression zone of the concrete by the soil, reducing the requirement for confinement reinforcement.
Fig.5-7: Confinement of compression zone by soil pressure for in-ground hinges (after Priestley and Seible, 1997)
To conclude with the integrity checks it is interesting to note that none of the countries that responded to the questionnaire mentions a requirement for a minimum reinforcement ratio for the piles