An investigation about seismic behavior of piled raft foundation for oil storage tanks using centrifuge modelling

Soil Dynamics and Earthquake Engineering 104 (2018) 210–227

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An investigation about seismic behavior of piled raft foundation for oil storage tanks using centrifuge modelling

MARK

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Seyed Mohammad Sadegh Sahraeiana,b, , Jiro Takemurab, Sakae Sekib a b

Department of Civil and Environmental Engineering, Shiraz University, Shiraz, Iran Department of Civil and Environmental Engineering, Tokyo Institute of Technology, Tokyo, Japan

A R T I C L E I N F O

A B S T R A C T

Keywords: Oil storage tank Liquefaction Piled raft foundation Centrifuge modelling

Some level of settlement is allowed in the design of oil tank if the uneven settlement can be controlled in an allowable value. Considering a critical condition of piled raft foundation (PRF), that is, secure contact of raft base to the ground surface, and the expected function of piles to impose additional resistance against the local settlement, PRF is considered as one of the rational foundation systems for the oil tanks. However, PRF has a complex interaction with soil under horizontal seismic loading, especially if the tank rests on a liquefiable soil, which may cause an extreme change of the soil stiffness underneath the tank. In this study, a series of centrifuge model tests was performed to investigate the mechanical behavior of oil tank supported by piled raft foundation on liquefiable saturated sand and non-liquefiable dry sand. In the tests, two types of foundation were modelled; a slab foundation, and a piled raft foundation. Using the observed results, such as accelerations of the tank and ground, dynamic and permanent displacement of the foundation, and excess pore water pressures of the ground, advantages and limitations of piled raft foundation for application to oil tanks on sandy soil are discussed.

1. Introduction

effective pile spacing [8,9]. Mechanical behavior of this foundation system under various loading conditions has been also studied by physical modelling. Static lateral loading tests were conducted in 1g condition to evaluate the application of pile groups and PRF, and discuss the optimized parameters, e.g. raft size, number of piles, piles spacing [10,11]. Furthermore, 1g experimental and analytical studies were performed for static lateral loading conditions to investigate the effects of pile head connection conditions between the raft and piles [12,13]. Similar researches were also made for dynamic loading conditions to investigate the performance of piled raft foundation [14,15]. In addition, some studies were accomplished about the performance of piled raft foundations, which experienced real earthquake loadings [9,16,17]. Centrifuge modelling is a prevalent approach for various studies in fields of geotechnical engineering, including soil liquefaction [18,19] and soil-structure-interaction problems [20,21]. To study the mechanical behavior of piled raft foundation as a complex soil-structure-interaction problem, centrifuge model tests have also been conducted under not only static loadings but also dynamic loadings. Horikoshi et al. [22] and Sawada and Takemura [23] used centrifuge modelling to compare the behavior of PRF with pile group and raft foundations under static horizontal loadings. On the other hand, Horikoshi et al. [14] and Nakai et al. [24] conducted dynamic centrifuge model tests to

Majority of existing oil storage tanks in Japan were constructed before the early 1970's when soil liquefaction was first considered in the design of tank foundations. Since the 1964 Niigata earthquake, the 1978 Miyagiken-oki earthquake [1] and the 1995 Hyogoken-Nambu Earthquake, it has become an urgent matter for geotechnical engineers to assess the seismic stability of existing oil storage tanks and implement proper countermeasures against soil liquefaction. Piled raft foundations (PRFs) have received considerable attention in the recent years, especially since Burland et al. [2] introduced the settlement reducer concept of PRF. The raft in this foundation system has adequate bearing capacity; therefore, the main objective of introducing the pile elements is to control or minimize the settlement, especially uneven settlement, rather than to carry the major portion of the vertical loads. Therefore, a major concern in the design of PRF is how to design the piles optimally to control the settlement [3–5]. Some researchers utilized finite element modelling (FEM) to study the effect of raft and pile dimension on the behavior of this foundation system [6,7]. Also, PRFs have been used for building design and some case studies of buildings have been reported. Field measurements were employed in these cases to estimate several parameters such as settlement, uneven settlement, load sharing between piles and the raft, and

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Corresponding author at: Department of Civil and Environmental Engineering, Shiraz University, Shiraz, Iran. E-mail addresses: [email protected], [email protected] (S.M.S. Sahraeian), [email protected] (J. Takemura), [email protected] (S. Seki).

http://dx.doi.org/10.1016/j.soildyn.2017.10.010 Received 2 March 2016; Received in revised form 2 August 2017; Accepted 16 October 2017 0267-7261/ © 2017 Elsevier Ltd. All rights reserved.

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sensors were placed in two different sections; one section at the center line of the model in the shaking direction; and the other, in the transverse direction. Model dimensions and instrumentation details are shown in Fig. 1(a), (b) and (c).

study the dynamic behavior of PRFs and pile group foundations, including the effect of pile head connection. Despite numerous studies on piled raft foundations, optimal and rational design methods of piled raft foundations have not been extended to the civil engineering infrastructures. This is partly due to the complex soil-structure interaction between raft, ground and piles during an earthquake. In particular, if the piled raft rests on a liquefiable ground, the soil-foundation interaction becomes more complex. Because of this complexity and possible large settlement, the practical implementation of piled raft foundation is further hindered. Another concern in the seismic design of piled raft foundation is to secure the contact of the raft with the subsoil. Without the contact, the contribution of raft cannot be assured against the horizontal load. To ensure the secure contact, the foundation settlement should be greater than the ground settlement. In the design of oil tank foundations, uneven settlement is a greater concern than maximum settlement. For example, an allowable uneven settlement is 1/300 of tank diameter [25], implying that some level of tank foundation settlement is permitted as long as uneven settlement is maintained below the allowable value. Therefore, piled raft foundation is considered one of the rational foundation systems for the oil storage tanks. Some studies have been conducted about oil tank foundations. For example, performances of pile foundation of storage tanks were investigated in some case studies [26,27]. Sento et al. [28] reported case studies about oil tanks on liquefiable sandy soil using compaction method as the countermeasure. A few researchers have considered piled raft foundations for the oil tanks. A case study of oil storage tank with piled raft foundation was done by Liew et al. [29]. Chaudhary [30] utilized FEM to study the behavior of piled raft foundation for a huge storage tank. As few researches on PRF of oil tanks on the liquefiable sand, Imamura et al. [31] and Takemura et al. [32] investigated about the dynamic response of oil tank supported by PRF using centrifuge modelling. From these researches, dynamic and permanent behavior of foundations were well observed. However, the observations were made in the shaking direction only, not in the different directions. In this study, dynamic centrifuge model tests were performed to investigate the mechanical behavior of oil tank supported by piled raft foundation on liquefiable saturated sand and non-liquefiable dry sand. In the tests, two types of foundations were modelled for oil storage tanks, namely, slab foundation (SF) and piled raft foundation (PRF). From the observed behavior, such as excess pore water pressures and accelerations of the ground, and accelerations, rotation and settlement of the tank, typical dynamic behavior and permanent displacements of the tank with PRF were studied and compared with those of the slab foundation not only in the shaking direction but also in the transverse direction. From these investigations and comparisons, the advantages and limitations of piled raft foundation for the application to oil tanks on sandy soil are discussed.

2.2. Tank, pile, raft and ground modelling Characteristics of the tank, pile and raft model used for the tests are presented in both the model and prototype scales in Table 2 (for more details about scaling factors in geotechnical centrifuge modelling refer to Garnier et al. [34]). The tank model (Fig. 2(a)) is made of an acrylic cylinder with 140 mm outer diameter, 160 mm height and 3 mm thickness. These dimensions were selected to model a small size tank considering the capacity of the model box. It was glued to the slab/raft model made of an aluminum disk with 150 mm diameter and 10 mm thickness (Fig. 2(a) and (c)). The raft model has 12 conical shape concave holes which are put onto the pile heads (Fig. 2(d) and Fig. 3). Silica sand No.8 (Table 3), which was used for the model ground, was glued to the bottom surface of the model raft to create a rough surface condition. Water was used as a liquid in the tank with a height of 140 mm. The total weight of the water, tank and raft (2.9 kg), created 1.42 kN of weight and 81 kPa of the average raft base pressure under 50 g centrifugal acceleration. The piled raft foundation has 12 identical piles, made of an aluminum tube with outer diameter of 6 mm, a thickness of 0.5 mm, and length of 100 mm as shown in Fig. 3. The rough piles shaft surface was also made by gluing silica sand No.8. These piles were arranged symmetrically as shown in Fig. 2(d). Utilizing this number and configuration of piles, the spacing/diameter ratio of piles (s/d) for most of the piles is 5.4. Friction angle of sand (φ’) with relatively medium condition (Dr = 65%) is about 40° [35]. The calculated vertical bearing capacities of the raft, assuming the full mobilization and partial mobilization (tan φ* = 2/3tan φ’ [36], φ* = 30°) of the friction angle of the sand, range from 29 to 147 MN and 18 to 92 MN for dry and saturated sand, respectively, in prototype scale. The vertical bearing capacity of one pile for these friction angles ranges from 0.32 to 1.7 MN and 0.19–1.1 MN for dry and saturated sand, respectively, in prototype scale. The total load of tank, including the tank and raft, is about 3.6 MN, which is much smaller than the bearing capacity of the raft alone, but almost larger than the total bearing capacities of the 12 piles. From these calculations, the expected function of piles as a settlement reducer which is the major objective of piles in PRF, can be confirmed. The pile heads were not rigidly fixed onto the raft, but simply capped by the concave hole, which allows free rotation like pinned connection (Fig. 3). In this way, the piles were subjected mostly to large axial and lateral forces and a small bending moment at the connection point to the raft. This condition is close to the actual situation of normal piledfoundation of oil tank [37]. In the model pile design, flexural rigidity and axial stiffness of concrete piles were targeted, but not failure of piles. As confirmed in Table 2, the axial load causing yield of the pile material is larger than the most of expected pile bearing capacity, shown above, and also much higher than the total load divided by pile number, i.e., 0.3 MN (= 3.6/12). The raft made by aluminum can be considered as a rigid plate which corresponds to a small diameter tank foundation. These conditions of structure components were made to focus on the effects of soil failure rather than the structural failures. In order to measure the pile axial load and shaft friction, the piles used in Case 2 were instrumented by axial strain gages at the head and tip as shown in Fig. 3. However, in Case 4, to prevent the non-uniformity of the ground made by sand pouring due to the wires connected to the piles, non-instrumented piles were substituted while 5 external (non-built-in) earth pressure cells (E.P.s) were glued on the raft base to measure the raft contact pressure (Fig. 2(d)). The raft model with nonbuilt-in E.P.s was also used in Case 3a. To improve the reliability of earth pressure measurement by eliminating the stress concentration on the attached E.P.s, a new raft model with 5 built-in E.P.s covered by

2. Dynamic centrifuge tests 2.1. Equipment, model foundations and test cases The centrifuge tests were conducted using Tokyo Tech Mark III centrifuge and a shaking table [33], under 50 g centrifugal acceleration. For modelling of the ground, a laminar box consisted of 15 laminas and rubber membrane bag with inner dimensions 600 mm in length, 250 mm in width and 438 mm in depth was used as shown in Fig. 1. Because the main objective in the current research was to model ground without liquefaction and with complete liquefaction, a simple uniform sandy ground with a moderate relative density was modelled beneath the tank. To this end, five model tests were performed (Table 1). In Case 1 and Case 2, a slab foundation (SF) and a piled raft foundation (PRF) were placed on dry sand, respectively. The SF and PRF were also modelled in Cases 3a and 3b and Case 4 for saturated sand. Case 3b was conducted in almost same conditions as Case 3a. The 211

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Fig. 1. Model setup, instrumentation and laminar box used for the tests.

thin silicon rubber were employed in Case 3b (Fig. 2(b)).

Table 2 Characteristics of tank, raft and pile used in model and prototype in 50 g.

2.3. Model preparation, and test procedure Tank

Fine silica sand (No.8) was used for the liquefiable sand layer and coarse silica sand (No.3) for the bottom drainage layer. Detailed properties of silica sands No.3 and No.8 are presented in Table 3. Dry silica sand No.3 was first placed at the bottom of the laminar box and then the fine silica sand layer was made by air pluviation. In the model, de-aired water was used as pore fluid of the sand. In order to ensure the ground saturation, CO2 gas firstly was passed from the bottom of laminar box through the dry sands slowly to replace the air with CO2 gas in the pores. The box was then placed in a vacuum tank. Having vacuum conditions less than − 95 kPa, de-aired water was introduced from bottom of the box [38]. The coarse grain size silica sand No.3 was utilized as the drainage layer at the model bottom to supply water evenly into the model ground during the saturation process. The sand layer with a relative density of 65% was aimed as target density, but in some cases, the final relative density had a few deviations from the target value (Table 1). The piles were fixed in the center of the modelling box by an aluminum guide during pouring the sand (Fig. 4). During the sand preparation, the accelerometers and pore water pressure transducers were placed at the prescribed locations as shown in Fig. 1. As water was used for pore fluid in the model, the prototype permeability of silica sand No.8 was about 1.0 × 10− 3 m/s in 50 g centrifugal acceleration (Table 3). Although this value is relatively high, it is low enough to accumulate excess pore water pressure and produce soil liquefaction in the early stage of the shaking. The permeability of soil affects the soil pile interaction, especially excess pore water migration in the vicinity of the pile. Gonzalez et al. [39] reported the effects of pore fluid (water and viscus fluid) on the displacement of piles subjected to lateral spreading by liquefaction. The difference of the pile displacement was attributed to extent of dissipation of suction which was generated by the dilatancy due to local shearing around the pile. The larger suction can be sustained for the less permeable case with

Raft

Pile

material outer diameter thickness height weight (liquid & raft) tank average pressure material diameter thickness base surface material outer diameter thickness length axial rigidity: EA yielding axial load bending rigidity: EI shaft surface

Model

Prototype

acrylic cylinder 140 mm 3 mm 160 mm 1.42 kN 81 kPa aluminum 150 mm 10 mm rough aluminum 6 mm (0.5 mm) 0.5 mm 10 mm 596 kN 0.60 kN 0.0023 kNm2 rough

steel 7.0 m 8.0 m 3.6 MN 81 kPa RC 7.5 m 0.5 m rough RC 0.3 m 25 mm 5m 1.49 GN 1.50 MN 14.2 MNm2 rough

viscous fluid and the relatively high prototype permeability of soil in the model might have some effect at pile vicinity. However, the suction around the pile is considered to be small because of less relative displacement between pile and soil in the model of this study, because of less lateral movement of soil and floating pile in the liquefied soil. Having made the model ground and placing the model tank on the ground, the tank was vertically loaded by an electrical jack to have a secure contact between the raft base and the ground surface. The loaddisplacement curves measured in the preloading for all cases except for Case 3a (the data were not recorded) are presented in Fig. 5. As shown in this figure, the maximum preload for different cases was chosen based on the sand stiffness and foundation type. In the cases with the piled raft foundation (PRF), it was assumed that the raft load proportion (RLP) will be about 50% and the preload about twice that of slab foundation (SF) was applied. Also, in saturated cases the preload was

Table 1 Test cases. Test code

Foundation

Ground

Details

Case Case Case Case Case

Slab Piled Raft Slab Slab Piled Raft

Dry sand (Dr = 66%) Dry sand (Dr = 68%) Saturated sand (Dr = 65%) Saturated sand (Dr = 68%) Saturated sand (Dr = 69%)

Slab w/o E.P.s 12 instrumented piles & raft w/o E.P.s Slab with 5 non-built-in E.P.s Slab with 5 built-in E.P.s 12 non-instrumented piles & raft with 5 non-built-in E.P.s

1 2 3a 3b 4

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Fig. 2. Tank and raft model. Fig. 3. Pile model.

Table 3 Properties of silica sands.

Specific gravity Mean grain size Effective grain size Uniformity coefficient Maximum void ratio Minimum void ratio Permeability coefficient(prototype for 50 g)

Gs D50 (mm) D10 (mm) Uc emax emin k(m/s)

No.8

No.3

2.65 0.1 0.041 2.93 1.333 0.703 2.0 × 10− 5 (1.0 × 10−

2.56 1.47 1.21 1.26 0.971 0.702 4.6 × 10− 3 (2.3 × 10− 1)

3)

half of the dry cases due to the looseness of the saturated sand in comparison to the dry sand. In this way, the maximum preloads on the dry sand were 490 N and 980 N for SF and PRF, respectively; while, in the saturated cases, about half of these loads were exerted on the foundations. The measured earth pressures in Case 3b with the built-in cells and Case 4 with the non-built-in cells during the preloading process are presented in Fig. 6, together with the average raft contact pressure exerted by the jack. In Case 4, the pressure could not be recorded by E.P.5 (the raft center). In Case 4, all the earth-pressure cells

Fig. 4. Guide and piles during sand pouring.

recorded larger values than the calculated average pressure (25 kPa), which was calculated neglecting the load supported by the piles. These larger stresses can be attributed to the stress concentration on the nonbuilt-in earth-pressure cells. However, the trends of variation in the measured contact pressures were well comparable to that of the average pressure, meaning that even those non-built-in cells could provide qualitative useful data during the test. In order to eliminate this 213

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than the others. In Shake 2, Case 3a and Case 3b had almost the same level of input motion, but they were remarkably larger than Case 4. From the time variation of Arias intensity, it can be seen that the majority of major input acceleration had been exerted until 10–15 s and thereafter, the input acceleration amplitudes were so small and the differences between Case 3a, 3b and 4 were negligible. In the shaking tests, the ground and tank accelerations, the horizontal and vertical displacements of the tank, and the excess pore water pressures in the ground were measured as shown in Fig. 1. In the saturated cases, these instrumentations were made not only in the longitudinal section of the shaking direction, but also in the transverse section (Fig. 1(c)). In the following discussions, the results of the model tests are given in the prototype scale unless otherwise stated. Fig. 5. Exerted loads in preloading process.

3. Results and discussions undesirable stress concentration, the built-in earth-pressure cells were implemented at the raft base in Case 3b. Although the measured contact pressures by the built-in cells showed a large difference, the average value of the measured pressures was much closer to the average exerted pressure. The inevitable uneven ground surface condition could be a reason for the large variation of the measured pressures. After this preloading process, the jack was detached and the whole setup was mounted on the shaking table on the swing platform of the centrifuge. Filling the tank with water, the centrifugal acceleration was increased up to 50 g. For the dry sand cases, white noise vibration was applied to the model, but not for the saturated sand cases to avoid the change of structure of sand. After confirming the steadiness of all sensor outputs, the shaking tests were conducted. The target input wave of the main shock used in the tests is the EW component of the acceleration recorded at Kurikoma, Kurihara city in 2008 Iwate-Miyagi Nairiku earthquake [40]. After the first shake, the second shake with about fifteen percent higher amplitude was applied to the model. The comparison of target acceleration and its Fourier spectrum with those of input motions in the tests are presented in the prototype scale in Fig. 7. Due to the limited performance of the shaker, high frequency component of the targeted motion could not be made. Furthermore, there were some differences in the input acceleration, which can be clearly seen in the variation of Arias intensity of the input accelerations in Fig. 8. Arias intensity (Ia) firstly proposed by Arias (1970) [41] is a measure of intensity of shaking defined as:

Ia =

π 2g

∫0

∞

[a (t )]2 dt

3.1. Tank on dry sand 3.1.1. Ground response Ground accelerations observed beneath the tank in Shake 1 for the SF and PRF models on dry sand are shown with the input accelerations in Fig. 9. Fourier amplitudes of the acceleration beneath the tank (A6) are compared to those of the input in Fig. 10. The ground accelerations were amplified towards the shallow depth in both cases, and they were amplified more in low periods (high frequencies) in Case 1 with SF. This behavior could be related to higher confinement pressure of the soil underneath the tank in SF than PRF (Case 2). As shown in subsequent section, in Case 2 the majority of tank load was supported by piles and small portion of the tank load was transferred to the soil beneath the tank; therefore, the stiffness of the soil under the tank with large confining pressure in Case 1 was larger than that in Case 2, resulting in the large predominant frequency of the subsoil. 3.1.2. Tank response, settlement and rotation Accelerations at the tank top and bottom in the shaking direction (A8, A9) during Shake 1 and their Fourier amplitude are shown in Fig. 11. The tank top acceleration was larger than that of the bottom in Case 1 with SF, implying a significant rocking motion. While in Case 2 with PRF, the difference between the top and bottom accelerations and Fourier amplitudes are not as much as Case 1 which show the efficiency of the piled raft foundation in reducing rocking motion of the tank. The spectra of accelerations at the tank bottom and just beneath the tank are almost the same in Case 1. While in Case 2, the acceleration of tank bottom is larger than that beneath the tank, implying that the shaking motion was transmitted through the piles and the tank vibrated rather independently from the ground. The raft contact condition to the ground caused the differences in the acceleration predominant period of the tank between the two foundations. The predominant period was smaller in Case 1 than Case 2 and the tank acceleration was largely

(1)

where a(t) is shaking acceleration and t is time. In the dry sand cases, almost the same input motion could be exerted for the two test cases in each shake, but in the saturated sand cases, the different motions were inputted between the test cases depending on the first and second shakes. In Shake 1, the input motions in Case 3b and Case 4 are nearly similar, especially until 7 s; but in Case 3a, it was significantly larger

Fig. 6. Raft base pressures during preloading in Cases 3b & 4.

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Fig. 7. Input accelerations and their Fourier spectra.

could be caused by the rocking motion enhanced during this period (Fig. 11) and the load transfer by the piles to deep stiff layer. The LDTs data show dynamic variation in both cases, but the amplitude of variations and its duration is partly larger in the case of slab foundation which is another evidence of large rocking motion in this case. Although the tank accelerations in the transverse direction were not measured in the dry sand cases, it can be confirmed from the small amplitude of L3 that the rocking should be smaller in the transverse direction than the shaking direction. In Fig. 14 the tank rotations in the two dry sand cases are shown. The rotation of the tank in the shaking direction (L1 - L2 direction) and L1 - L3 direction are presented together with the maximum rotation for Case 1 while in Case 2 due to the absence of L2 data the rotation is presented only in L1-L3 direction. Considerable difference between the two foundations can be also seen in the rotational behavior. Both the rotation amplitude in dynamic response and the residual rotation of the slab foundation were larger than those of the piled raft foundation, which verify a better performance of PRF in reducing the tank rotation during dynamic loading. The rotation of SF ceased at t = 8 s and remained constant, despite the further shaking at the later part of the shaking. The same trend can be confirmed in PRF, but the time of becoming the constant rotation angle was earlier than SF. It should be noted that the rotation in the shaking direction was much smaller than that in L1-L3 direction. The rotational behavior will be discussed more

amplified in this small period range (0.3–0.5 s). This tendency of tank response was also observed in the white noise vibration as shown in Fig. 12. But the predominant period of the white noise vibration was shorter than that in the shaking test, suggesting the reduction of soil stiffness by the large input motion in the latter than the former. Tank settlements measured by three laser displacement transducers (L1, L2 and L3, Fig. 1) are compared for the slab and piled raft foundations in Fig. 13. In Fig. 13(a) the settlements measured at the opposite top edges of the tank in the shaking direction (L1, L2) are shown, while in Fig. 13(b) the tank edge settlements in the transverse direction (L3) are shown with the tank center settlement, that is, the average of L1 and L2. The tank center settlement and L2 are not plotted in the figures for Case 2 as the settlement could not be measured by L2 in this case due to the dislocation of the laser from the target plate. However, it is confirmed from the figures that the settlements of PRF are much smaller than those of SF, which is a good evidence of settlement reducer function of the piles. In the beginning of shaking until 2.5 s, although the input accelerations were not so large, both cases showed rapid settlements, which could be attributed to the densification of sand. After the initial settlement, the settlement rate once decreased in both cases, but in Case 1 the rate increased again from around 6–8 s and then the tank showed very small settlement increment until the end of shaking. This large increase of settlement is not clearly seen for Case 2 in the figure. The difference in the trend of settlement from 6 to 8 s

Fig. 8. Arias intensity of input motions.

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Fig. 9. Ground accelerations beneath the tank in Shake 1 (Dry sand).

the shaking, implying larger contribution of the pile resistance against the rotational moment from the tank as illustrated in Fig. 15(c). Total loads carried by all piles are presented in Fig. 15(b). In the figure, the measured pile head resistance, that is, the tank load carried by the piles, the pile tip, and shaft resistance components are shown together with the tank load. It should be noted that due to the interference of the moment strain to the axial strain measurement near the pile top, the measured total pile load was overestimated, which could be seen in the static and dynamic components. Namely, the measured pile loads were

in later section. 3.1.3. Piles behavior Using the instrumented piles in Case 2, the head axial load, tip resistance and shaft friction carried by each pile were measured. Fig. 15(a) shows the axial forces time histories of piles 1, 5 and 9 (Fig. 2(d)). The pile at the inner part of the raft (No. 9) carried larger bearing load in the static condition. On the other hand, the pile at outer edge in the shaking direction (No. 1) showed larger amplitude during

Fig. 10. Ground responses in Shake 1 (Dry sand).

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Fig. 11. Tank responses in Shake 1 (Dry sand).

Fig. 12. Ground and tank responses during White noise (Dry sand).

for this particular PRF. About 70% of pile resistance was mobilized by the tips and 30% by the shaft in the static condition. In order to verify the effects of piles and raft on the rotational behavior of the tank, the time-histories of the tank rotational moment, piles resistant moment and raft resistant moment

larger than the tank load in static conditions and during the shaking, though the shaking motion was applied in the horizontal direction, not vertical direction. However, the pile axial load bearing behavior can be discussed by the measured data and it can be said from the figure that the majority of the tank load was carried by piles in the static conditions

Fig. 13. Tank settlements in Shake 1 (Dry sand).

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Fig. 14. Tank rotations in Shake 1 (Dry sand).

Considering the contribution of piles against static and dynamic loads in the dry sand conditions, more rational design of PRF could be made by reducing the number of piles. However, for reducing the pile number, the other critical conditions rather than the foundation settlement, such as failure of the piles and punching of the raft should be assessed for the design of PRF for oil tanks on dry sand, in which the relatively large axial pile load could be generated with the static and dynamic component.

are shown in Fig. 15(d) and the tank rotational moment and the piles resistant moment are compared in Fig. 15(e). The tank rotational moment and piles resistant moment are calculated from the tank inertia force and the piles axial forces respectively around the center of the raft (Fig. 15(c)). The difference between these two moments is the raft resistant moment. Although the error in the measurement of pile axial forces could cause the uncertainty in the estimation of the moment resistance, the mobilization trend of the resistance can be confirmed in Fig. 15(d) and (e). Both of the tip and shaft resistances almost evenly contributed in preventing the tank rotation (Fig. 15(e)). The piles resistant moment had the main role in bearing the tank rotational moment. Despite very small raft contact pressure in the static condition, it is confirmed that the raft also resisted the moment load.

3.2. Tank on saturated sand 3.2.1. Ground response Fig. 16 presents the ground responses at shallow depth (Z =

Fig. 15. Piles behavior in Case 2 (PRF, Dry sand). (a) Variation of piles loads, (b) Total axial force carried by piles, (c) Diagram of piles resistant moment (PRM) and tank rotational moment (TRM), (d) Tank rotational moment, piles and raft resistant moment during the shaking, (e) Piles resistant moment vs. tank rotational moment.

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Fig. 16. Ground accelerations in Shake 1 (saturated sand).

becomes long due to the liquefaction, which could cause the amplification in the long period range as seen beneath the tank in Case 3b and Case 4. In Case 3a, because of relatively large input motion, the attenuation was more than the other two cases. The acceleration response of A11 located at 0.75 m out from the raft edge in the transverse section was similar to that of A5 beneath the tank center, implying that the tank load could affect the dynamic behavior at the location of A11. In the results of Cases 3b and 4, which had almost similar input motions, a significant difference cannot be confirmed between SF and PRF in the saturated sand while it was observed in the dry sand.

1.25 m) in the saturated cases (Case 3a, 3b and 4) during Shake 1. The accelerations recorded beneath the tank (A5: X = 0 m), and beside the tank (A6: X = 6.75 m, A7: X = 10 m) along the center line of the tank in the shaking direction (Y = 0 m), and at the transverse section (A11: X = 0 and Y = 4.5 m) are compared with the input motion (A1). Fourier amplitudes of the accelerations measured by A5, A7 and A11 are compared with that of input motion in Fig. 17. Attenuation of input motion was observed in the ground, especially for the short period component, which is a clear evidence of soil stiffness reduction by liquefaction. The attenuation was more significant beside the tank than beneath the tank, which can be attributed to the confinement effect of the tank to the soil underneath. This prevents significant reduction of stiffness and increase of damping ratio as happened in the free ground beside the tank. However, even though the natural period of the ground

3.2.2. Excess pore water pressure The excess pore water pressures (EPWP) in the saturated sand cases (Cases 3a, 3b and 4) at different locations of the ground observed in 219

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Fig. 17. Ground responses in Shake 1 (saturated sand).

Shake 1 are presented in Fig. 18. The figure shows the EPWP in different depths in four vertical arrays, three in the longitudinal section at the tank center (P7, P4, P3 and P2: X = 0 m), near (P8 and P5: X = 6.75 m) and far (P9 and P6: X = 10 m) outside of the raft, and one in the transverse section (P12, P11: X = 0, Y = 4.5 m). In Case 3b, P2 and P3 could not be recorded. In the figure, the initial overburden stress ′ = zγ ′, γ ′: effective unit weight of sand) and a nominal vertical stress (σv0 ′ and the stress by the tank (σv′ * ) are presented. σv′ * is the sum of σv0 pressure, which is calculated by the elastic solution assuming the uniformly distributed raft pressure on an elastic half-space and RLP = 100% [42]. Although this assumption (RLP = 100%) is not completely correct, according to later discussions (Section 3.2.3), during the liquefaction period, most of the tank load is carried by the raft; therefore, this assumption is not so unrealistic during the shakings. The σv′ * value is a criterion for discussion on the behaviors in different cases and σv′ * – EPWP is a representative value for the remaining effective stress in the ground during the liquefaction. In the early stage of shaking, the EPWPs increased rapidly and then this rapid rise ceased and the water pressure either became almost constant or increased gradually during the shaking. Then, the EPWPs started the dissipation. At the locations, outside of the raft (P8, P5, P9 and P6) where the tank load did not affect σv′ * remarkably, the EPWP built nearly up to the σv′ * value, meaning almost zero effective stress. While at the location beneath and near the tank (P7, P4, P3, P2, P12 and P11), where the tank load affected σv′ *, the EPWPs were smaller than the σv′ * values, reconfirming the confinement effect of the tank on the soil underneath. The larger the dif′ is (more confinement effect from the tank), ference between σv′ * and σv0 the larger the remaining effective stresses, that is, σv′ * – EPWP. Although the effective stress in the area beneath the tank didn’t reach zero value

as compared to the surrounding area, the effective stress is affected considerably due to the large EPWP beneath the tank. The pore water pressure behaviors of Cases 3b and 4 were almost the same, except P7 and P4 where EPWP in the rapid increase was larger in Case 4 than Case 3b while Case 3a showed different behavior especially in the late start of the dissipation. Furthermore, the residual EPWPs observed at the shallow depth after dissipation was larger in Case 3a than the others. These residual EPWPs were due to the tank settlement at the location beneath the tank and the settlement of PPT due to the relatively large unit weight at the location out of the raft. The possible reason of the difference between Case 3a and the other cases is the differences in the liquefaction level due to the larger input motion and smaller relative density in Case 3a than the others (See Figs. 7 and 8 and Table 1). The relatively large EPWP in the early stage of shaking beneath the tank in PRF (Case 4) than SF (Case 3b) could be attributed to the increase of raft pressure due to the reduction of pile bearing load. As above mentioned, the typical variation of EPWPs can be divided into three parts. The first part is until the end of rapid increase, which is called “build-up period” and (t1), is the end of this period. The second part is from (t1) to the time when the pore pressure starts the dissipation (t2), which is named “liquefaction period”. The third is the dissipation stage from (t2) until the end of this period (t3). Determining the end of dissipation (t3) is not straightforward due to the gradual decrease of EPWP and the residual EPWP. The end of dissipation was determined at the time when the decrease of EPWP from the maximum value became 99% of that at the end of the measurement. These three times (t1, t2 and t3) are highlighted in the EPWP graphs of Fig. 18 and are used for the discussion on the later parts.

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Fig. 18. Excess pore water pressures of the ground in Shake 1(saturated sand).

were more uniform than those measured in Case 3b with the built-in cells, which was partly because of the effect of initial surface undulation on the measured pressures which was more significant for the latter cell type than the former. However, comparing the variations of Case 3a and 3b, two points can be confirmed. The first is that the built-in cells

3.2.3. Raft base contact pressures Fig. 19 shows the variations of raft base contact pressures measured by five pressure cells during Shake 1. In Case 4, the cell at the tank center (EP5) and in Case 3a, EP3 could not measure the pressure. The measured base pressures in Cases 3a and 4 with the non-built-in cells

Fig. 19. Raft base pressures of Cases 3a, 3b and 4 in Shake 1(saturated sand).

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Fig. 20. Average raft base pressures in Case 4 (PRF, saturated sand).

significant in Case 3a and Case 3b than Case 4. In the transverse direction, the accelerations were much smaller than those in the shaking direction. In particular, the bottom transverse acceleration is almost negligible and with no clear dominant frequency. But the top transverse accelerations showed a certain vibration with a clear dominant period of about 0.3 s, which is about half of the dominant period in the acceleration of the shaking direction, 0.6 s. This clear response of the tank top in the transverse direction can be attributed to the deflection of tank top due to the relatively small hoop stiffness at the top part of the tank wall.

could measure the dynamic pressure better than the non-built-in cells and the second is that the average value of the five pressures measured by the built-in cells during the shaking was closer to the average tank pressure acting on the raft (σT = 81 kPa) while the non-built-in cells all recorded the pressure more than σT . From these facts, it can be said that the built-in cells could give more precise pressure measurements than the non-built-in ones. The base pressures in Case 3b showed very interesting behavior that the pressure at the center (EP5) increased, while the others at the outer part of the raft decreased, implying the stress concentration at the center portion during the shaking. It could be attributed to the less reduction of stiffness at the tank center than the perimeter side as discussed in Fig. 18. Due to the failure in measuring EP5, this stress concentration could not be confirmed in the piled raft model (Case 4). However, the different behavior in Case 4 can be seen from that in Case 3b. All pressures at the raft outer side showed a quick rise at the EPWP generation period (t1) and then started decreasing around the time t2 of P7. The measured average raft base pressure by the four outer pressure cells in Shake 1 and Shake 2 are shown with EPWP at P7 in Fig. 20. Much clearer trends can be confirmed from the figure, that is, the raft load proportion increased by the reduction of pile loads due to the liquefaction, but the raft load decreased, in other words, the pile load was regained gradually by the recovery of effective stresses of the soil due to the dissipation of EPWPs. The recovery of pile load was earlier in Shake 2 than Shake 1, which corresponds to the fact that t2 in Shake 2 was earlier than that in Shake 1 for Case 4 (Fig. 20). The average raft base pressure at the outer side of the raft after the shaking was larger for Shake 1 than Shake 2. There are two possible reasons for that; one is the larger pile load proportion for Shake 2 than Shake 1, and another one is more stress concentration on the center part for Shake 2 than Shake 1. It should be noted that although the variation of the raft base pressures is not shown for Shake 2, the more stress concentration on the center part was confirmed after Shake 2 than Shake 1 from the fact that the residual center pressure in Case 3b after Shake 2 was about 270 kPa larger than that after Shake 1 (220 kPa).

3.2.4.2. Tank settlements. The tank settlements were measured by the three laser displacement transducers at the locations shown in Fig. 23. In Shake 2 of Case 3a, as the LDTs could not record the data for certain time intervals, three sets of the settlement could be measured only until t = 4 s and the measurements of L1 and L3 could be resumed near the end of the test. In Fig. 23 the settlements at the tank center, which is the average of L1 and L2 are compared for the entire period and the early stage of shaking at the top and bottom figures with t2 and t1 of P7 and P8 respectively. The settlements increased gradually during the shaking in contrast to EPWPs behavior, that is, a quick rise in a short time (Fig. 18). Comparing the results in Shake 1 and Shake 2, the effect of densification by the first shake can be confirmed. Even though the input motion in Shake 2 was larger than that of Shake 1, the settlements in the second shake were much smaller than those of the first, except for Case 3b of which the second input motion was significantly larger than the first one, about 2 times in terms of Arias intensity (Fig. 8). The effect of pre-shaking was more significantly evidenced in the beginning. In Shake 1, the tank started settling at time of about 1.5 s, which is the actual onset of the shaking in terms of Arias intensity (Fig. 8), and the settlement rate increased with time until t = 3 s while in Shake 2, there were no substantial settlement until t = 2.3 s and no increase in settlement rate was observed in the beginning. The relatively large settlement of Shake 1 in the beginning of shaking could be attributed to the poor contact of the raft base to the ground surface, a kind of bedding error, which can be removed by the first shake. After this initial part, the settlement increased almost linearly with time until t = 8 s at the time when EPWP dissipation started in the deep depth beneath the tank (P2 as shown in Fig. 18). Although the settlement rate decreased at this point, further settlement occurred even after the t2P7, the time when EPWP just beneath the raft (P7) started decreasing, until t2P8, the time when EPWP at the shallow depth beside the raft (P8) started decreasing. After t2P8, the minor settlement, which was mainly caused by small shaking and the consolidation of sand, took place. The settlement after t2P8 seems smaller for Case 4 (PRF) than Cases 3a and 3b (SF). Recovery of pile bearing load, which can be confirmed in Fig. 20 could be a reason for the smaller settlement in the late stage of the shaking and after the shaking in the PRF than the SFs. In the period of liquefaction, from t1 to t2, the settlements of slab foundations were larger than that of the piled raft foundation. However, as the larger settlements of SF could be partially due to the larger input motions (Fig. 8), the effectiveness of PRF as a settlement reducer against the soil liquefaction could not be confirmed in these tests.

3.2.4. Tank response 3.2.4.1. Acceleration. The accelerations at the top and bottom of the tank in the shaking direction (A9, A8) are shown with the input acceleration in Fig. 21. In the figure, the tank response in the transverse direction is also shown. The Fourier amplitude spectra of the acceleration both in the shaking and transverse directions are indicated in Fig. 22. In the shaking direction during the rapid increase of EPWP in the build-up period before t1, difference of the input, and the tank bottom and top accelerations were relatively small but after t1 the bottom and top accelerations showed difference, bigger at the top and smaller at the bottom than the input, resulting rocking motion of the tank. The rocking motion in Case 3a with the large input motion was higher than both other cases, but Case 3b has less rocking motion in comparison with Case 4 even though the input motion in this case was almost same until t = 6 s and slightly larger after that time than Case 4 (Fig. 8). In all the cases, short period components were significantly attenuated but the long period components amplified especially in Case 3a and Case 4. As a difference between SF and PRF, it can be seen that the phase difference between the tank bottom and top accelerations were more 222

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Fig. 21. Tank response accelerations in Shake 1 (saturated sand).

figure, results of Case 2 are missed, because L2 could not be measured. Also, for Shake 2 of Case 3a, limited results are shown, as the three settlements could be obtained until t = 4 s and L1 and L3 could give the data at the end of the shaking. In the beginning of shaking, the rotation gradually increased with time in all the cases. The increase of rotation, in the beginning, is earlier and much larger for Shake 1 and Shake 2 of Case 1 (SF, Dry sand) and Shake 2 of Case 3a than the other cases. The large tank accelerations of dry sand from the beginning (Fig. 11) can be a reason for this behavior. As discussed in Fig. 14, the rotation of SF on dry sand ceased at t = 8 s and was kept constant and the same trend was observed in Case 2 (PRF, Dry sand). The tank accelerations measured in the beginning of Shake 2 are depicted for the saturated cases in Fig. 25. A large amplification was

3.3. Tank maximum rotation (dry and saturated cases) In the stability assessment of tank foundation, the uneven settlement is a critical concern. For the relatively small diameter tank supported by a rigid slab or raft, the uneven settlement is equivalent to the rotation of the foundation. In the previous dynamic model tests of foundation using the one directional shaking table, e.g., Takemura et al. [32], the rotation of tank foundation was only measured in the shaking direction. In this study, with the settlement at three locations (Fig. 23), the maximum rotation and its direction were measured. Fig. 24 shows the revolution of the maximum rotation during the shaking for the entire period and the early stage of shaking in the top and bottom figures with t2 and t1 of P7 and P8, respectively. In the 223

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Fig. 22. Tank response Fourier spectrum in Shake 1 (saturated sand).

these results, the effect of PRF in preventing uneven settlement cannot be confirmed against the liquefaction in the entire depth of piles. However, in the beginning of the shaking in the period of EPWP buildup, the rotations of PRF were smaller than those of SF, showing the effectiveness of PRF for the condition that the pile vertical bearing resistance could be mobilized. Fig. 26 shows revolutions of the direction of maximum rotation, θ, during the shaking. The definition of θ is given in the figure. As there was some inevitable error in the estimation of direction from the measured settlement, especially in the beginning of shaking when the settlement was very small, therefore the first part of the data with large

observed at the tank top from the very beginning in Case 3a as compared to the two other cases which can be the reason for the large rotation in Shake 2 of Case 3a. On the other hand, the behavior of the tank on saturated sand is quite complicated and varied in different cases. In Case 4 (PRF), the rotation increased monotonically in the liquefaction stage (from t1 to t2) both for Shake 1 and Shake 2. While in Shake 1 of Cases 3a and 3b, the rotation increased after t1 but decreased in the liquefaction period. In Shake 2 of Case 3b, the rotation behavior fluctuated more during the shaking. As a result, the tank rotations after the shaking was larger for Case 4 than Cases 3a and 3b, except for Shake 2 of Case 3a, which showed the large rotation from the beginning. From

Fig. 23. Tank center settlements (Saturated sand), top: entire shaking period with t2, bottom: earlier stage of shaking with t1.

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Fig. 24. Tank maximum rotations, (a) entire shaking with t2 (b) earlier stage of shaking with t1.

fluctuation is eliminated in the figure. The observed behavior in the direction of maximum rotation can be divided into two groups. In the first group, the direction suddenly changed from the shaking direction (θ = 0 or 180) to the different direction and became constant to a certain direction, which was observed in the both shakes of Case 1, Case 4 and Shake 2 of Case 3a. While in the second group, the direction gradually changed from the shaking direction until the end of the shaking, which was observed in Shake 1 of Case 3a and both shakes of Case 3b. By comparing Fig. 26 with Fig. 24, it can be said that the large rotation occurred once the direction was almost fixed in the first group. While in the second group, the rotation did not monotonically increase, but decreased during liquefaction stage between t1 and t2. Even in Shake 2 of Case 3b, the rotation, and the direction both showed fluctuation. From these observations, it can be inferred that once the direction of the rotation is fixed to a direction diverted from the shaking direction, the rotation will be accumulated by shaking. But while the rotation mainly takes place in the shaking direction in the beginning of shaking, the monotonic increase of the rotation may not easily occur and the rotation behavior becomes very complicated as seen in the slab foundation cases (Fig. 24). However, it should be noted that even for the slab foundation, once a relatively large tilting to the diverse direction was triggered by a disturbance, such as a large rocking motion of the tank, a large rotation occurred inevitably by the shaking as shown in Shake 2 of Case 3a. There could be several reasons why Case 4 (PRF) tilted in the diverse direction in the early stage of shaking, such as: inevitable difference of bearing resistance of each pile, non-uniformity of raft base contact condition to the ground, which could cause the rotation of the tank to the area with small pile resistances and poor contact of the raft base to the ground.

The relationships between the tank maximum rotation and the tank center settlement were presented in Fig. 27. In the top figure, the relationship obtained from the dry ground cases are shown with the results of the previous study [32] about the relationship between the rotation of shaking direction and the tank settlement. While in the bottom of Fig. 27, the same relationship was plotted for the saturated cases with the indication at the time of EPWP buildup (t1) and liquefaction stage (t2) obtained from the location of P8 at the shallow depth beside the tank. Due to the lack of settlement data, the rotations of L1L3 direction are used for some cases, as shown in the figures, but in Shake 2 of Case 2, even this data could not be recorded. In dry cases, the trend of relation is almost the same for SF and PRF, except the PRF of the previous study, which shows much smaller uneven settlement than the others. This small rotation could be considered as an advantage of PRF. However, the advantage may be overestimated because of the uncertainty of the rotation in the diverse direction and the non-monotonic increase of the rotation. Nonetheless, the effectiveness of PRF for preventing the uneven settlement can be confirmed in the dry sand cases (Case 1 and Case 2). Namely, the tank settlement caused by the shaking can be reduced by PRF, which results in the smaller rotation as compared to SF. In the saturated cases, the relationships are very different for the different cases and between the first and second shakings. As an overall trend of the relations, the following can be pointed out: 1) majority of the settlement and rotation took place in the liquefaction stage, 2) the rotations relative to the settlements were larger for the second shake than the first shake, 3) in the early stage until t1, the rotations of PRF were smaller than or equal to those of SF. From the limited results, the effectiveness of PRF in reducing uneven settlement cannot be confirmed especially for the condition of liquefaction of the

Fig. 25. Tank response accelerations in the beginning of Shake 2 (saturated sand).

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Fig. 26. Directions of tank maximum rotation.

Fig. 27. Tank maximum rotations vs settlements.

just underneath the tank, which also affects the EPWP behavior of the sand. 5. During the EPWP build-up period, as the pile bearing resistance can still be mobilized, the PRF can have effectiveness in reducing the tank uneven settlement. But once the liquefaction occurs in the entire depth of piles, the PRF cannot secure the effectiveness. Therefore, the effectiveness of PRF in reducing the tank uneven settlement on the partially liquefied ground or the ground with nonliquefied soil layers should be studied to clarify the conditions in which the PRF can be positively applied as a foundation of oil tanks. 6. From the observations of the maximum rotation of the tank, it can be confirmed that once the direction of the rotation is fixed to an orientation diverted from the shaking direction, the rotation will be accumulated by shaking. But while the rotation mainly takes place in the shaking direction, the monotonic increase in the rotation may not easily occur and the rotation behavior becomes very complicated. The direction of maximum settlement can be caused by inevitable difference of bearing resistance of each pile, non-uniformity of raft base contact condition to the ground. 7. Due to the pile load caused by the static and dynamic loadings in PRF on dry sand, special consideration is necessary for assessment of pile failure and punching failure of the raft, especially for the PRF with small number of piles. While for the case of soil liquefaction, the critical conditions are settlement and uneven settlement. For the actual implementation of the concept of PRF in the design of tank foundation, further study is necessary covering the various factors, especially level of liquefaction such as liquefaction potential (PL value) and duration of shaking.

entire depth of piles. Therefore, the effectiveness of PRF on the ground with partial liquefaction, such as, the liquefaction in partial depth, not entire depth, should be examined to clarify the condition for that PRF can be positively applied as a foundation of oil tanks. 4. Conclusions In this study, a series of dynamic centrifuge model tests has been conducted on a slab foundation (SF) and a piled raft foundation (PRF) of oil storage tank resting on dry and saturated sand. The following conclusions were derived from the centrifuge model tests. 1. In dry sand ground, the input motions were amplified in both types of foundation, but the amplification for the PRF was more in longer period range than the SF. The less confinement effect of the raft stress of PRF than that of SF made the natural period of the subsoil longer for the PRF than the SF. 2. In dry sand, not only the raft resistance but also the pile resistance bears the loads and rotational moment. Due to the contribution of piles of the pile raft foundation, the PRF can effectively reduce tank rocking motion, settlement and uneven settlement of the tank. 3. In the saturated sand, due to the significant stiffness reduction by liquefaction, the input motion was attenuated in the short period range, especially beside the tank with less confinement pressure from the tank. While in the long period range, the natural period of ground was increased and the input motion could be amplified. 4. In the saturated sand, the excess pore water pressure (EPWP) behavior of the sand underneath the tank is affected by the variation of raft contact stress during the shaking. Due to the decrease of effective stress caused by the liquefaction, pile resistance of PRF significantly decreases and the raft bearing load increases, which causes the rapid and relatively larger EPWP increase underneath the tank than that for SF. While in the liquefaction period, due to the recovery of pile resistance from the deeper depth, the raft bearing load starts decreasing earlier than the decrease of EPWP of the sand

References [1] Ishihara K, Kawase Y, Nakajima M. Liquefaction characteristics of sand deposits at an oil tank site during the 1978 Miyagiken-oki earthquake. Soils Found 1980;20(2):97–111. [2] J. B. Burland, B. B. Broms, and V. F. B. DeMello, Behaviour of foundations and structures, In: Proceedings of the 9th ICSMFE, 1977, vol. 2, pp. 496–546.

226

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S.M.S. Sahraeian et al. [3] Poulos HG. Piled raft foundations: design and applications. Géotechnique 2001;51(2):95–113. [4] Poulos HG, Small JC, Chow H. Piled raft foundations for tall building. Geotech Eng J SEAGS AGSSEA 2011;42(2):78–84. [5] Horikoshi K, Randolph MF. A contribution to optimum design of piled rafts. Géotechnique 1998;48(3):301–17. [6] Ziaie-Moayed R, Kamalzare M, Safavian M. Evaluation of piled raft foundations behavior with different dimensions of piles. J Appl Sci 2010. [7] A. Alnuiam, H. El Naggar, and M. H. El Naggar, Performance of Piled-Raft System under Axial Load, In: Proceedings of the 18th Inter Con Soil Mech. Geotech. Eng., 2013, p. 2663–66. [8] Yamashita K, Yamada T, Hamada J. Investigation of settlement and load sharing on piled rafts by monitoring full-scale structures. Soils Found 2011;51(3):513–32. [9] K. Yamashita, Field measurements on piled raft foundations in Japan, In: Proceedings of the 9th Inter. Conference on Testing and Design Methods for Deep Foundations (IS-Kanazawa 2012), Japan, Kanazawa, 2012, vol. 1, pp. 79–94. [10] K. Pastsakorn, Y. Hashizume, T. Matsumoto, Lateral load tests on model pile groups and piled raft foundations in sand, In: Proceedings of the International Conference on Physical Modelling in Geotechnique Geotech, 2002, 709–714. [11] J. Hamada, T. Tsuchiya, T. Tanikawa, and K. Yamashita, Lateral loading model tests on piled rafts and their evaluation with simplified theoretical equations, In: Proceedings of the 9th Inter. Conference on Testing and Design Methods for Deep Foundations (IS-Kanazawa 2012), Japan, Kanazawa, 2012, vol. 1, p. 467–76. [12] Matsumoto T, Fukumura K, Pastsakorn K, Horikoshi K, Oki A. Experimental and analytical study on behaviour of model piled rafts in sand subjected to horizontal and moment loading. Inter J Phys Model Geotech 2004;4(3):1–19. [13] Matsumoto T, Nemoto H, Mikami H, Yaegashi K, Arai T, Kitiyodom P. Load tests of piled raft models with different pile head connection conditions and their analyses. Soils Found 2010;50(1):63–81. [14] Horikoshi K, Matsumoto T, Hashizume Y, Watanabe T. Performance of piled raft foundations subjected to dynamic loading. Int J Phys Model Geotech 2003;3(2):51–62. [15] Matsumoto T, Fukumura K, Horikoshi K, Oki A. Shaking table tests on model piled rafts in sand considering influence of superstructures. Inter J Phys Model Geotech 2004;4(3):21–38. [16] Yamashita K, Hamada J, Onimaru S, Higashino M. Seismic behavior of piled raft with ground improvement supporting a base-isolated building on soft ground in Tokyo. Soils Found 2012;52(5):1000–15. [17] Yamashita K, Hashiba T, Ito H, Tanikawa T. Performance of piled Raft Foundation subjected to Strong seismic motion. Geotech Eng J SEAGS AGSSEA 2014;45(2). [18] R. Dobry and L. Liu, Centrifuge modelling of soil liquefaction, In: Proceedings of the Tenth World Conference on Earthquake Engineering, Madrid, Spain, 1992, p. 6801–9. [19] Dashti S, Bray JD, Pestana JM, Riemer M, Wilson D. Centrifuge testing to evaluate and mitigate liquefaction-induced building settlement mechanisms. J Geotech Geoenviron Eng 2010;136(7):918–29. [20] Dashti S, Hashash YMA, Gillis K, Musgrove M, Walker M. Development of dynamic centrifuge models of underground structures near tall buildings. Soil Dyn Earthq Eng . 2016;86:89–105. [21] Olarte J, Paramasivam B, Dashti S, Liel A, Zannin J. Centrifuge modeling of mitigation-soil-foundation-structure interaction on liquefiable ground. Soil Dyn Earthq Eng . 2017;97:304–23. [22] Horikoshi K, Matsumoto T, Hashizume Y, Watanabe T, Fukuyama H. Performance

[23] [24]

[25] [26]

[27]

[28]

[29]

[30] [31]

[32]

[33]

[34] [35]

[36] [37] [38]

[39]

[40] [41] [42]

227

of piled raft foundations subjected to static horizontal loads. Int J Phys Model Geotech 2003;3(2):37–50. Sawada K, Takemura J. Centrifuge model tests on piled raft foundation in sand subjected to lateral and moment loads. Soils Found 2014;54(2):126–40. S. Nakai, H. Kato, R. Ishida, H. Mano, and M. Nagata, Load Bearing Mechanism of Piled Raft Foundation during Earthquake, In: Proceedings of the Third UJNR Workshop Soil-Struct. Interact., 2004. Notification of technical specifications regarding to the regulations on hazardous materials. FDMA (Fire Disaster Management Agency, Japan), 1974. M. Cubrinovski and K. Ishihara, Analysis of the performance of an oil-tank pile foundation in liquefied deposits, XV ICSMGE TC4 Satell. In: Proceedings of the Conference Lessons Learn. Recent Strong Earthq., 2001. B. H. Fellenius, Tech, and M. Ochoa, Large liquid storage tanks on piled foundation, In: Proceedings of the Inter Conference Found. Soft Ground Eng.-Chall. Mekong Delta, 2013, p. 3–17. N. Sento, S. Yasuda, N. Yoshida, and K. Harada, Case studies for oil tank on liquefiable sandy ground subjected to extremely large earthquakes and countermeasure effects by compaction, In: Proceedings of the 13th World Conference Earthq. Eng., 2004. S. Liew, S. Gue, and Y. Tan, Design and instrumentation results of a reinforcement concrete piled raft supporting 2500 t oil storage tank on very soft alluvium deposits, In: Proceedings of the 9th Inter Conference Piling Deep Found., 2002. Chaudhary MTA. FEM modelling of a large piled raft for settlement control in weak rock. Eng Struct 2007;29(11):2901–7. S. Imamura, T. Yagi, and J. Takemura, Dynamic stability of oil tank supported by piled-raft foundation on liquefiable sand, In: Proceedings of the Physical Modelling in Geotechnics, Zurich, 2010, vol. 2, p. 1409–14. J. Takemura, M. Yamada, and S. Seki, Dynamic response and settlement behavior of piled raft foundation of oil storage tank, In: Proceedings of the Physical Modelling in Geotechnique, Perth, Australia, 2014, vol. 1, p. 613–9. J. Takemura, M. Yamada, and S. Sakae, Dynamic response and settlement behavior of piled raft foundation of oil storage tank, In: Proceedings of the Joint 9th Inter. Conference of CUEE/4th Asia Conference on Earthquake Eng., Tokyo, 2012, p. 533–44. Garnier J, et al. Catalogue of scaling laws and similitude questions in geotechnical centrifuge modelling. Int J Phys Model Geotech . 2007;7(3):01–23. Yamada M. A study on dynamic response and seismic settlement behavior of piled raft foundation for oil storage tank (M.Sc. Dissertation). Tokyo, Japan: Tokyo Institute of Technology; 2012. Terzaghi K. Theoretical soil mechanics. New York, USA: John Wiley & Sons, Inc; 1943. Ishimatsu S, Yagi T, Yoshimi T, Takemura J. Filed observation of pile behavior during the liquid level variation in an oil tank. Kisoko 2009;37(10):76–9. T. Kimura, J. Takemura, A. Hiro-oka, M. Okamura, and T. Matsuda, Countermeasures against liquefaction of sand deposits with structures, In: Proceedings of the IS-TOKYO’ 95, Tokyo, Japan, 1995, vol. 3, pp. 1203–1224. Gonzalez L, Abdoun T, Dobry R. Effect of soil permeability on centrifuge modeling of pile response to lateral spreading. J Geotech Geoenviron Eng . 2009;135(1):62–73. JMA, Japan Meteorological Agency, 2008 〈http://www.seisvol.kishou.go.jp/eq/ kyoshin/jishin/080614_iwate-miyagi/index.html〉. Kramer SL. Geotechnical earthquake engineering. Pearson Education; 1996. B. M. Das, Fundamentals of Geotechnical Engineering, 3rd edition. Chris Carson.