京 都 大 学 防 災 研 究 所 年 報 第 50 号 C 平 成 19 年 4 月
Annuals of Disas. Prev. Res. Inst., Kyoto Univ., No. 50 C, 2007

Advanced Techniques for Evaluating Vulnerability of Urban
Infrastructure byIntegrating Multiple Evaluation Indexes

Masayoshi NAKASHIMA, Hitoshi TANAKA, Tomotaka IWATA,
Sumio SAWADA, and Shuji TAMURA
Synopsis
It is essential for disaster prevention to quantify the earthquake vulnerability of urban
infrastructures in consideration of their life cycle. As part of a five-year study on this subject,presented in this article are: the refinement of prediction of strong motions, examinations
into actual performance of nonstructural components, rational evaluation of RC building
structures in consideration of foundation-structure interaction, effects of total earth pressure
on pile forces, and development of a systematic procedure for the estimation of life-cycle
cost for large-scale lifeline network. Development of a sloshing reduction damper which
consists of separation wall with slits installed in cylindrical tank.
Keywords: strong ground motion, performance of nonstructural components,
foundation-structure interaction, earth pressure, sloshing reduction damper

1. Introduction
The rational evaluation of earthquake vulnerability of
our urban infrastructures require accurate prediction ofstrong motions, thorough characterization of theperformance of structures, soils, structure-building
system, and soil-structure systems, and development ofmethodologies that can effectively consider thelife-cycle cost/performance of the structures andsystems. This is a five-year project being undertaken
within the scope of Center of Excellence for Natural
Disaster Science and Disaster Reduction, implemented
by the Disaster Prevention Research Institute, Kyoto
University. Achievement of fiscal 2006 and five-years
is summarized in what follows.
2. Strong Ground Motion Prediction for Design
Ground Motion
2.1 Long-period ground motion prediction during
the hypothetical Tonankai earthquake
At 19:07 and 23:57 on 5 September 2004, an

MJMA7.1 and a MJMA7.4 earthquakes occurred in
the south east off the Kii peninsula, Japan, that werenamed the 2004 Kii-Hanto-Nanto-Oki earthquake.
These earthquakes and their aftershocks are thought tobe intraplate earthquakes, which occurred in the
subducting Philippine Sea Plate (The Headquarters
for Earthquake Research Promotion, hereafter
HERP, 2004), that was neighbor of source region of
the hypothetical Tonankai earthquake. In order toestimate ground motions during the hypothetical
Tonankai earthquake, we used one of aftershock
events using the empirical Green’s function (EGF)
method by Irikura (1986). First, we determine thesource size and the stress drop of the EGF event
using comparison of the aftershock and thepre-event records. Those parameters were
determined as 28.1 km2 and 3.3 Mpa (Suzuki et al.,2005). Using this event as the EGF again, we
anticipated the hypothetical Tonankai earthquakeand simulated the resultant strong ground motions.
Although the type of earthquake was different, thefocal mechanism of the EGF event acceptably－9－resembled that of the hypothetical Tonankai
earthquake. We followed Model 2 used in HERP
(2001b). This model consists of three asperities andgives a larger stress drop on the smaller asperities, as
shown in Fig. 1. We only characterized the asperity
portion, and ignored the background off-asperity area.
The rupture starts from off the Kii peninsula andpropagates in a north-easterly direction. We calculated
the acceleration, velocity, and displacement
waveforms predicted at the stations K-NET, KiK-net,
and CEORKA, where the records of the EGF event
were available. The frequency band used for thissimulation was 0.05–20.0Hz. The rupture velocity was
assumed to be 2.7 km/s, referring to HERP (2001b).
Figure 2 shows the distribution of seismic intensity(JMA, 1996) calculated from simulated acceleration
waveforms. This is consistent with that calculated by
HERP (2001b), which used the stochastic Green’s
function method. Large seismic intensities of morethan 5+ were predicted for the Aichi and Mie
prefectures, which is located toward the rupture
propagation. Examples of the velocity waveforms forthe stations in the Aichi and Mie prefectures (AIC016
and MIE003) and the Osaka basin (FKS and AMA)
are shown in Fig. 3. At the stations such as FKS and
AMA, that is in the Osaka basin, long-period groundmotions last for more than 3min. Although seismic
intensity in the Osaka basin is up to 5+, such long
period about 3 to 6 second ground motions will hit the
Osaka basin site during the Tonankai earthquake.
Predominant period of predicted ground motions at thestations in the Osaka basin is different at the sites andthat indicates the complex three-dimensional basin
underground structure affects such the ground motion
characteristics. Especially, the sites at the Osaka bay
area have longer period ground motions and there aremany long period structures such as high rise buildings
and fuel tanks. It is necessary to consider long period
ground motions during the large subduction
earthquakes to reduce the seismic disaster against long
period seismic waves.
At MIE003, long-duration ground motions are also
observed after the direct S-wave. On the other hand,
the waveform predicted at AIC016 mainly contains ahigh frequency component with very large amplitude.
Seismic intensity 7 was predicted for this station. The
depth to the bedrock at MIE003 is deeper than that at
AIC016, although both stations are located in the Nobibasin. That is also indicated that the complexity of theunderground structure of the Nobi basin. Constructionof the large basin underground velocity structure areneeded.

Fig.1 Source model of the hypothetical Tonankai
earthquake used in this study. Each star indicates the
rupture starting point of each asperity.

Fig. 2:Distribution of the calculated seismic
intensity.

Fig. 3. Example of horizontal velocity waveforms
predicted for the stations in the Aichi and Mie
prefectures (AIC016 and MIE003) and in the
Osaka basin (AMA and FKS).
2.2 Ground motion modeling of the 2005
Miyagi-Oki earthquake
We estimate the source model of the 2005
Miyagi-Oki earthquake, which consists of two strong－ 10 －

motion generation areas (SMGAs). The location of thefirst SMGA is determined from the arrival timedifferences between the first motion and the first pulsein the P-wave portion. Based on the broadband
waveform modeling, we estimate the two SMGA
parameters by using the empirical Green’s function
method. The estimated location of the two SMGAs
coincides with the two major large slip areas inferredfrom the inversion of strong motion and teleseismic
waveforms as shown in Fig. 4. It implies that thebroadband strong motion, including the high-frequency
range of the 2005 Miyagi-Oki earthquake, is mainly
radiated from the asperities. The deeper SMGA has alarger stress drop (34.1 MPa) than the shallower
SMGA (17.6 MPa).
In Fig, 5, we compare the SMGAs of the 2005 eventand those of the 1978 event estimated by Kamae etal.(2002). Even using the newly-determined
hypocenters by Okada et al.(2005), SMGAs of both
events do not overlap each other.

Fig. 5. Comparison of the SMGA location of the 2005
Miyagi-Oki earthquake with that of the 1978
earthquake estimated by Kamae (2006). The solid
squares are the SMGAs of the 2005 event, and thedotted rectangles are those of the 2002 event. (a) The
hypocenters determined by the JMA are used as
reference points. (b) The distribution is shifted toadjust the hypocenters to those determined by Okada
et al. (2005).
3 Performance of Nonstructural Components inEarthquake Conditions
3.1 Importance of nonstructural components inseismic design
Most research on the behavior of building structuresduring an earthquake only focuses on the lateral force
resisting system. However, nonstructural elements also
play a critical role in the performance of a structure asdamage to these systems can result in significant
economic losses or loss of life. In order to address thisissue, this study characterizes the behavior of both
nonstructural partition walls and hanging ceilingsystems using the shake table at the Disaster
Prevention Research Institute of Kyoto University. The
test setup allowed for the interaction between
perpendicular walls and the hanging ceiling system tobe considered.

Fig. 4. The location of the estimated SMGAs projectedon the slip distribution inferred by Wu and Koketsu(2005) by using strong-motion and teleseismic data,
together with the strong motion stations using
waveform modeling. The slip distribution is shifted tomatch the epicenter used in their study to that
determined by the JMA or Hyp0.

3.2 Test setup and specimens
In order to perform the shake table study of thenonstructural partition walls and hanging ceilingsystems, a steel frame is designed to provide a meansof imparting inertial loads to the system with the massplaced at the top of the loading frame. The loading
frame is designed such that it will have minimal effecton the behavior of the non-structural system. Fig. 6(a)provides a view of the loading frame.
Two nonstructural specimens are tested as part ofthis study. Both specimens have the same nonstructural－ 11 －

partition wall design with the difference between them
being the design of the hanging ceiling system(traditional and seismic). The nonstructural partition
walls are typical gypsum board and light gage steel
construction used for Japanese high-rise officestructures. Further details in regards to the constructionof these systems can be found in Lee et al. (2007). The
partition walls are constructed on four sides so that theeffects associated with perpendicular walls can bestudied. The walls span 3.35 m in the long direction(X-direction) and 1.35 m in the short direction(Y-direction). In the X-direction, a door is framed into
one of the walls, while the other wall has a 1.5 m wide
opening. The two short direction walls are solid. The
partition walls can be seen in Fig. 6(b).

Fig. 7 Traditional hanging ceilingFig. 8 Seismic hanging ceiling
The systems are loaded using a sinusoidal motionthat increases in frequency during its duration. The
magnitude of the acceleration of the sinusoidal motion
is determined so as to apply maximum drifts of 1/200,
1/100, 1/66, 1/50 (or 1/35), and 1/28 radians with eachsubsequent run. At each drift level, the motion isapplied in the X-direction and Y-direction separately.

(a)

(b)
Fig. 6 Test Setup: (a) Bare loading frame; (b)nonstructural specimen.
The traditional hanging ceiling uses evenly spaced
threaded rods and clips to hold the ceiling, and theceiling itself abuts up against the partition walls. The
seismic ceiling uses light gage steel angles to form atruss system to further stabilize the ceiling. A gap isalso left between the ceiling and the partition walls.
The seismic ceiling is designed for drift levels of 1/100
radians. Figs. 7 and 8 contain pictures of the twoceiling systems.

3.3 Observed damage to system
Damage to the nonstructural partition walls andhanging ceiling systems are observed after each drift
level to correlate the damage with drift. During loading,the nonstructural partition walls undergo a rocking
behavior as a result of each of the presence of
perpendicular walls. This rocking behavior results inearly damage occurring at the corner of the walls. At
1/100 radians, a gap between the door frame andgypsum board formed. At a drift of 1/66 radians thedoor became difficult to open. Larger drift levels of
1/50 radians resulted in the screws pulling through the
gypsum board minimizing the attachment to the metal
studs. Cracks also began to propagate in the gypsum
board from the upper corners of the doorway. Some of
the gypsum board panels become detached at drift
levels of 1/28 radians.
The traditional hanging ceiling systems performed
well at all drift levels. The seismic hanging ceilingsystem also performed well up to its design drift of
1/100 radians. However, at drift levels of 1/66 radians,
pounding between the ceiling and walls began to occur
resulting in some of the clips between the truss system－ 12 －

and ceiling being dislodged. All of these clips are lostat drift levels of 1/28 radians reducing the ceiling to amore traditional configuration, but the ceiling itself
remained basically undamaged.

60

Shear Force (kN)

40
20
0
-20
-40

3.4 Partition wall behavior
The behavior of the nonstructural partition walls is
evaluated using the shear force sustained by thenon-structural system versus drift curves when loadedin the X-direction. In general, both specimens showedsimilar behavior. The hysteretic curves for the 1/200,
1/100, and 1/66 radian drift levels are shown in Fig. 9.
The hysteretic curves show some general trends
including nonlinearity of the hysteresis, stiffness andstrength degradation with continued cycling, andpinching of the hysteresis with increasing drift levels.
The shear force reaches a maximum value of 46 kN at
the 1/35 radian drift level.
The stiffness and damping associated with this
hysteretic behavior is also considered. From Fig. 9, it
can be seen that the initial stiffness decreases with anincrease in drift level as the system sustains damage.
The secant stiffness also shows a similar decrease from1.33 kN/mm to 0.23 kN/mm between the 1/200 radianand 1/28 radian drift cases. The change in energy
dissipation based on the area enclosed by the hysteretic
curve follows a similar trend as the shear force withthe maximum value of 3349 kN-mm occurring during
the 1/35 radian drift case.
30

Shear Force (kN)

20
10
0
-10
-20
-30
-0.005-0.002500.0025

0.005

Inter-story Drift Angle (rad)(a)
45

Shear Force (kN)

30
15
0
-15

-60
-0.02

-0.01

0

0.01

0.02

Inter-story Drift Angle (rad)(c)
Fig. 9. Shear versus drift behavior for the nonstructural
partition walls: (a) 1/200 drfit; (b) 1/100 drfit; and(c) 1/66 drift.
3.5 Hanging ceiling system behavior
The traditional and seismic hanging ceiling systems
are evaluated by comparing the measured ceilingacceleration to the acceleration at the top of the loading
frame. For both types of ceiling systems, the values
remained similar up to drift levels of 1/66 radians,
since no significant damage is sustained. At the drift
level of 1/28 radians, the traditional ceiling system has
accelerations 25% below that of the loading frame,
while the seismic ceiling systems’ accelerations are 1.5
times that of the loading frame. The difference can be
attributed to the gap associated with the seismic ceiling,which resulted in pounding between the ceiling andwall. The maximum accelerations sustained by the
traditional ceiling system and seismic ceiling system
are 1.3g (1/35 radians) and 2.9g (1/28 radians),
respectively.
3.6 Conclusion
1) Nonstructural partition wall systems containing
perpendicular walls tend to be dominated by arocking behavior during loading.
2) The door system undergoes permanent damage
and decreased functionality at drift levels of 1/66
radians and drift levels of 1/28 radians cause
falling debris from the wall systems.
3) Both ceiling systems sustain minimal damage up
to their intended design level with the traditional
system resultinginsmallermaximum
accelerations.

-30
-45
-0.01

-0.0050

0.005Inter-story Drift Angle (rad)0.01

4 Experimental
mechanisms

Study

on

seismic

resisting(b)
4.1 Lateral Force Resisting Mechanism of a Multi
㧙Story Shear Wall and Peripheral Members－ 13 －

(1) Introduction
In current design procedures [RC design
code:AIJ:1999] [Paulay1992], cantilever structural
walls are normally assumed to stand on a solid
foundation, and foundation beams, slabs and piles aredesigned separately without considering theirinteractions. This study aims to experimentally clarify
the variation of the lateral load resisting mechanisms
considering the interaction between a shear wall, afoundation beam, slabs and piles, and to establish morerational design procedures for each structuralcomponent. The experimental parts of this study
were already submitted to this report last year. With
macro model of this specimen, the restoring force
characteristic of the shear wall was predicted anddamage evaluation of the foundation beam under themulti-story shear wall was conducted. This analytical
study was performed by Masanobu SAKASHITA,
Susumu KONO, Hitoshi TANAKA and Fumio
WATANABE.

simulate loading conditions of the prototype
fourteen-story shear wall system under earthquakes. At
the roller support, 0.7Q was applied horizontally to thepile on the compressive side and 0.3Q was applied tothe piles on the tensile side by a 1000kN jack in the
opposite direction to the 1000kN horizontal jack.

(2) Specimens
The specimen configuration was determined fromtypical fourteen-story residential buildings in Japan.
The assemblage consisting of the lowest three floors ofa shear wall with a foundation beam, the first floor slab,
and two piles in the transverse direction was scaled to
1/4. Figure 10 shows specimen configuration.

(4) Macro Model
Figure 12 shows macro model of this specimen.
Shear wall was modeled based on Suzuki et al‘s study
in 1993 and Miki et al’s study in 2002. Columns andbeams were modeled with rigid bodies, axial and shear
springs. The angle of the concrete strut was fixed to be
45 degrees. Because this specimen failed in flexure,these struts were modeled to be elastic and not able totransfer tensile axial force. As to longitudinal bars ofcolumn, beam and foundation beam, the effect of thetension stiffening was considered using the model of
Maekawa et al. (1999).

㪣㫆㪸 㪻㫀㫅 㪾㩷㪙 㪼㪸 㫄㪙㪼㪸㫄㪪㪿㪼㪸㫉㩷㫎㪸㫃㫃㩿㪮㫀㪻㫋㪿㩷㪎㪇㫄㫄㪀㪣㪸㫋㪼㫉㪸㫃㩷㪚㫆㫃㫌㫄㫅㪙㪼㪸㫄㪈㪝㩷㪝㫃㫆㫆㫉㩷㪪㫃㪸㪹㩿㪫㪿㫀㪺㫂㫅㪼㫊㫊㩷㪎㪇㫄㫄㪀㪫㫉㪸㫅㫊㫍㪼㫉㫊㪼㩷㪝㫆㫌㫅㪻㪸㫋㫀㫆㫅㪙㪼㪸㫄㪪㫃㪸㪹㩷㪮㫀㪻㫋㪿㩷㪈㪌㪇㪇㫄㫄㪝㫆 㫌㫅 㪻㪸 㫋㫀㫆 㫅㩷㪙㪫 㪿 㪼㩷㪺㪼 㫅 㫋㪼 㫉㩷㪻 㫀㫊 㫋㪸 㫅㪺 㪼㩷㪹 㪼 㫋㫎㪼㪸 㫄㪼 㪼 㫅㩷㫇 㫀㫃㪼 㫊㩷㩷㪊 㪇㪇㪇㫄㫄㪫㪿㪼㩷㪿㪼㫀㪾㪿㫋㩷㫆㪽㩷㪸㩷㫇㫆㫀㫅㫋㩷㫆㪽㩷㪺㫆㫅㫋㫉㪸㪽㫃㪼㫏㫌㫉㪼㩿㫋㪿㪼㩷㪺㪼㫅㫋㪼㫉㩷㫆㪽㩷㫋㪿㪼㩷㫇㫀㫅㪀㪈㪉㪌㪇㫄㫄㩷
㩷㩷

Fig.10 Specimen configuration(3) Test Setup
As shown in Figure 11, lateral load, Q, was applied
statically through a 1000kN horizontal jack to theloading beam. Two 2000kN vertical jacks were
adjusted to create appropriate column axial forces, N1
and N2, which are a liner function of lateral load, Q, toFig. 11 Loading system(5) Restoring Force Characteristic and Deformation
Shape
Figure 13 shows Lateral load – 1F Drift angle
relationship of the shear wall. Analyses were
performed by two different methods. One method was
conducted with shear wall-foundation model shown in
Figure 12. The other method was conducted with shear
wall model. This model has the same shear wall as
shear wall-foundation model, but doesn’t have thefoundation beam and the base of the shear wall was
fixed to the ground.
Analytical results of shear
wall-foundation model agreed with the experimental
envelope curve.
Figure 14 shows deformation shape of shear－ 14 －

wall-foundation model when drift angle was 0.30%.
Due to rotation of the compressive pile involving thetransverse foundation beam, deformation of thecompressive column of the shear wall was supposed tobe smaller than that in case of using fixed shear wall
base.
Nc

Nt
Q

Loading Beam

Axia Springl

3F Shear WallShear Spring

Figure 15 (a) (b) show strain distribution of
longitudinal bars of foundation beam when flexural
drift angle = 0.046% and 0.20%. Transition of thedamage of the foundation beam was confirmed inexperiment according to the flexural deformation of
the shear wall above the foundation beam. This model
simulates well the transition of the strain distribution of
the upper and lower longitudinal bars of the foundation
beam.
0.2

Strain of longitudinal bars (%)

2F Shear Wall
Beam
Column1F Shear Wall
Transverse
FoundationBeam㪬 pper㩷Longitudinal
Bars ofFoundation Beam
Lower
Longitudinal
Bars ofFoundation Beam

Pile
Pin

0  7Q

Pin roller

㪯 Location of 㪈㪌㪇㪇㫄㫄

㪇㫄㫄㪈㪌㪇㪇㫄㫄

0.15

Experiment (Upper)
Experiment (Lower)
Analysis (Upper)
Analysis (Lower)

0.1

0.05

0

-0.05
-1500 -1000 -50005001000 1500

Location of foundation beam (mm)Foundation Beam

Fig. 12 Macro model

Drift Angle, Flexural Drift Angle, Lateral LoadExperiment (0.132%, 0.046%, 320.6kN)
Analysis (0.168%, 0.046%, 314.6kN)(a) Relationship 1

300

0.25

200

Strain of longitudinal bars (%)

㪣㪸㫋㪼㫉㪸㫃㩷㪣㫆㪸㪻㩷㩿㫂㪥㪀400㪜㫏㫇㪼㫉㫀㫄㪼㫅㫋㪘㫅㪸㫃㫐㫊㫀㫊
㪘㫅㪸㫃㫐㫊㫀㫊㩷㩿㪦㫅㫃㫐㩷㫊㪿㪼㪸㫉㩷㫎㪸㫃㫃㪀㩷

100

0

0

0.2

0.4
0.6
0.8
㪛㫉㫀㪽㫋㩷㪘㫅㪾㫃㪼㩷㩿㩼㪀1

Fig. 13 Lateral load – first story drift relation

Experiment (Upper)
Experiment (Lower)
Analysis (Upper)
Analysis (Lower)

0.2
0.15
0.1
0.05
0
-0.05
-0.1
-1500 -1000 -50005001000 1500

Location of foundation beam (mm)Drift Angle, Flexural Drift Angle, Lateral LoadExperiment (0.637%, 0.203%, 354.8kN)
Analysis (0.815%, 0.203%, 361.1kN)(b) Relationship 2

Fig. 15 Strain distribution of longitudinal barsof foundation beamCompressive
Pile

Fig. 14 Deformation shape of macro model(Drift Angle= 0.30%)(6) Strain distribution of longitudinal bars of foundation
beam(7) Conclusions
One 1/4-scale cantilever structural wall system wastested and comparison was conducted withexperimental results and analytical results based on
strut-and-tie theory. The main conclusions can besummarized as follows.
1) The envelope curve of the load-drift angle－ 15 －

viscosity of 40 times that of water, was used.
A sketch and conditions of a piles-footingsuperstructure model is shown respectively in Fig. 17
and Table 1. A 2 × 2 pile model was used for the tests.
The lid-shaped footing was modeled with rigid brass.
The pile heads were linked rigidly to the upper plate of
the footing; their tips were also rigidly linked to thelaminar shear box. The strain gauges at the pile heads
were not in contact with the soil. Therefore, the shear
force at the pile head, as calculated by the strain
differentiation, was inferred to have been measured
accurately, and independently of the soil effects. Plates
supported by three load cells were set up on the rightand left sides of the footing to evaluate earth pressure
acting on the footing. The surface of the footing model
is smoother than that of a prototype. Therefore, a finegrade of sandpaper #800 was pasted on the right andleft sides and sidewalls of the footing. The
superstructure was modeled with rigid brass. The mass

relationship was simulated well with the macro model
including the foundation.
2) Transition of the strain distributions of thelongitudinal bars of the foundation beam in experiment
was simulated well.
4.2 Effects of Earth Pressure and Sidewall Friction
on Pile Stress during Soil Liquefaction(1) Objectives
This study is intended to: 1) examine the effects ofliquefied soil layer on development of earth pressure
and sidewall friction acting on an embedded footing in
non-liquefied crust; and 2) investigate the effects of
total earth pressure and sidewall friction on the shear
force at pile heads. For those purposes, dynamic
centrifuge tests on a superstructure-footing model thatis supported on piles in either dry sand or in saturatedliquefiable sand were performed.

$FFHOHURPHWHU
Accelometer
Silica sand N o.2 

Silica san d N o.2

Pore pressuretransducer

AccelerometerSilica san d N o .2

Strain gage

Plate
Footing

Load cellS ilica sa nd No .2

Su perstru cture

D ry san dD r= 45%

Sa turate d sa nd
D r= 45 %  চ 㧕(a) Case 1

Pile(b) Case 2

Fig. 16 Setup for centrifuge tests

Fig. 17 Pile-footing-superstructure model(2) Model Preparation
Centrifuge tests were performed at 40 g centrifugal
acceleration using the geotechnical centrifuge at theDisaster Prevention Research Institute, Kyoto
University. A pile-footing-superstructure model was
prepared in a laminar shear box with inner dimensionsof 450 mm (length) × 150 mm (width) × 200 mm(height). Two cases of the performed tests arepresented in Fig. 16. The soil profile in Case 1 consistsof a 4.0 cm layer of dry coarse sand crust, which isToyo silica sand #2 (D50=1.92 mm) overlying a 16.0
cm layer of dry fine sand deposit, which is Albany
silica sand #48 (D50=0.29 mm). The soil profile inCase 2 resembles that of Case 1, except that the finesand deposit is saturated and the water level was
located at G.L. -3.0 cm. A viscous fluid made of
Metolose (Shin-Etsu Chemical Co.), which has aof the superstructure was 1.035 kg; that of the footing
was 0.715 kg. The natural frequency of thesuperstructure under the fixed footing conditions was
about 100 Hz. All tests were excited using Rinkai92,
which is a synthesized ground motion for the Tokyo
Bay area. All data presented in the following sections
are of prototype scale.
(3) Effects of Earth Pressure and Sidewall Friction on
Shear Force at Pile Head
Figures 18 and 19 show the time histories of thesuperstructure-footing inertia defined by the sum of thesuperstructure inertia and footing inertia, and the sumof shear force at the pile heads. The shear force at thepile heads is similar in waveform to thesuperstructure-footing inertia in the non-liquefaction
test. Contrastively, the shear force differs greatly in－ 16 －

waveform from the superstructure-footing inertia in theliquefaction test. The shear force amplitude increases
600
Force(kN)(a) Superstructure-footing inertia

0that of superstructure-footing inertia. The amplitude offooting displacement tends to be greater than that ofsoil displacement, suggesting that the total earthpressure and sidewall friction act on the footing asreaction forces.

Force(kN)

-600
600

Table 1 Conditions of piles-footing-superstructure(b) Shear force at pile headmodel in prototype and model scale

0

-600
0

10

20

30
Time (s)

40

50

60

3LOH
)RRWLQJ

Fig.18 Time histories of superstructure-footing
inertia and shear force at pile head in Case 1

Force(kN)600

6WUXFWXUH(a) Superstructure-footing inertia

0-600
Force(kN)600(b) Shear force at pile head
0

-600
0

10

20

30
Time (s)

40

50

60

Fig. 19 Time histories of superstructure-footing
inertia and shear force at pile head in Case 1
rapidly after time t=10 s. It is noteworthy that the shear
force amplitude in the liquefaction test is greater thanthat in the non-liquefaction test, although thesuperstructure-footing inertia amplitude in theliquefaction test is less than that in the non-liquefaction
test.
To elucidate the difference in the shear force
between Cases 1 and 2, Fig. 5 shows the time historiesof the superstructure-footing inertia, the sum of thetotal earth pressure and sidewall friction, shear force atthe pile heads, soil displacement, and footing
displacement for time t=20–30 s. In the
non-liquefaction test, the sum of the total earthpressure and sidewall friction tends to be out of phaseby 180 degrees with the superstructure-footing inertia
andhas
smaller
amplitude
thanthesuperstructure-footing inertia. Such tendencies indicatethat the sum of the total earth pressure and sidewall
friction, which are generated as reaction forces of thesuperstructure-footing inertia, counter the inertial force
transmitted from the superstructure-footing to the pilehead. Therefore, the shear force amplitude is less than

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In Case 2, the amplitude of the sum of the total earthpressure and sidewall friction tends to be greater thanthat of superstructure-footing inertia. The shear force
resembles the sum of the total earth pressure andsidewall friction in waveform, which indicates that theshear force is caused mainly by the total earth pressure
and sidewall friction. The amplitude of soil
displacement tends to be greater than that of footing
displacement, suggesting that the total earth pressure
and sidewall friction act on the footing as external
forces.
(4) Conclusion
The liquefied layer under the non-liquefied crust
affects not only the superstructure-footing response but
also the earth pressure and sidewall friction. The
amplitudes of the total earth pressure and sidewall
friction in liquefaction test are much greater than thosein non-liquefaction test. In addition, the total earthpressure and sidewall friction in liquefaction test tend
to be in phase with the superstructure-footing inertia.
Consequently, the shear force amplitude at the pileheads becomes large. In non-liquefaction test, the total
earth pressure and sidewall friction, generated asreaction forces of the superstructure-footing inertia,
counter the inertial force transmitted from thesuperstructure-footing to the pile head. Therefore, theamplitude of shear force is smaller than that ofsuperstructure-footing inertia.

－ 17 －

Hz (3.2 Hz in some cases), whose displacement ismanually controlled to keep 1 mm. The height ofliquid surface is measured near outer wall ofcylindrical model tank, and recorded for 15 sec with
200 Hz sampling frequency. Note that effect offloating roof is not considered.

Force(kN)

600(a) Case 1
0Superstructure-footingᘠᕈജinertiaǛߖࠎᢿജ
Shearforce at pile headDis.(cm)

-600
2

Total
ᘠᕈജ㧙ߖࠎᢿജ
earth pressure + Sidewall fricton(b) Case 1
0

-2
20

Groundsurface displacement
#5##%%A&#(6#%%A&
Footing
displacement

21

22

23

24

25
Time (s)

26

27

28

29

30

Force(kN)600
ᘠᕈജSuperstructure-footing
inertiaᘠᕈജ㧙ߖࠎᢿജ
Total
earth pressure + Sidewall friction

0(c) Case 2

Ǜߖࠎᢿജ
Shesr
force at pile headDis.(cm)

-600
30

0
Ground surface displacement
Footing displacement
#5##%%A&(d) Case 2
-30
20

21

#(6#%%A&

22

23

24

25
Time (s)

26

27

28

29

30

Fig.20 Time histories of superstructure-footing
inertia, sum of total earth pressure and sidewall
friction, and shear force at pile head in Cases 1
and 2

Photo. 1 Shaking Table Experiment

Stainless steel frame

5. Development of Sloshing Reduction Dumper for
Petroleum Tank with Floating RoofGap sensor

5.1 Introduction
During 2003 Tokachi-oki earthquake, large
amplitude of sloshing occurred in many petroleum
tanks, and caused severe damage and fire. We will
have the similar damage during Tokai, Tonakai andNankai earthquake which are predicted to occur within
30- 50years. To reduce sloshing for petroleum tankswith floating roofs we propose a damper which
consists of a separating wall with slits installed in thetank. We perform shaking table experiments to
examine its effect on reduction of sloshing.
5.2 Model
The model used in the experiment is made of acryl
cylinder whose diameter is 512mm, height 310mm, as
shown in Fig. 21 and Photo 1. White colored water isused for measuring its surface by a laser gap sensor.
The separation wall, which consists of stainless steel
frame and brass panels, is installed in the model tankas shown in Fig. 21. 21 cases of panel layout, shownin Fig. 22, are examined in this study. The black area
indicate the brass panels and the white area is open.
The open ratio, D, is used to represent characteristics ofthe panel layout. Input motion is one directional
sinusoidal wave with the frequencies from 1 Hz to 3

Brass plate
LoadcellInput motion

Fig. 21 Experimental Model

Fig. 22 Panel arrangement experimented
5.3 Effect on Reduction of Sloshing－ 18 －

Fig. 23 compares the non-dimensional frequency
response function of sloshing height, that is divided by
the displacement of input sinusoidal wave, in case #00
and case#07 which are the case with no separation wall
(D=100%) and full separation wall (D=0%). In thecase #00 we can recognize clear peaks at 1.2 Hz and2.2 Hz which represent 1st mode and 2nd mode natural
frequencies. Note that the peak value of 1st mode
reaches 50 and that of 2nd mode is 20. It should be
the reason why large sloshing occurs during
earthquakes. In case #07, the peak at 1st mode
disappears and new two peaks at 1.7Hz and 2.5Hz
appear. These new peaks are generated because thetank is completely separated into two half parts and 1st
mode and 2nd mode for the half part are excited. We
call those modes “1.5th mode” and “2.5th mode” hearafter.

To reduce the sloshing height during earthquakes,
response values at 1.5th mode and 2.5th mode are also
important as well as 1st mode. Fig. 25 shows the
response function for cases #11 and #06 whish havethe same 25% open areas at different position on theseparation wall. The figure implies that open area inthe center of the wall reduces the sloshing height at
1.5th mode.

Fig. 25 Response Function ( #11 and #06)

Fig. 23 Response Function (#00 and #07)
Fig. 24 shows the open ratio of the panel layout on
the separation wall affects the response function. We
find that small open ratio results small sloshing heightat 1st mode and enlarge the 1.5th mode. On the other
hand 2nd mode and 2.5th mode are not significantly
changed.

From the experimental results of other cases shownin Fig.21 we find that the layout with many small open
area reduces the sloshing height at 2nd and 2.5th mode.
We finally find that case #16 is the best case in the 21
cases experimented, which have small sloshing height
through the wide frequency range. Fig. 26 shows the
response function of case #16 in comparison with case#00, that is the case of no separation wall. Sloshingheight at 1st mode reduced from 50 to 10. the height at
2nd mode also decreased from 20 to 10. On the other
hand, 1.5th mode and 2.5th mode are newly excited
while these heights are controlled to be about 10.

Fig. 26 Response function for the best case ( #16 ) andno separation wall case (#00)

Fig. 24 Response Function ( #01, #04 and #06)

5.4 Conclusion
We propose a damper which consist of separationwall with slit installed in cylindrical tank in order toreduce the sloshing height during earthquake. Model－ 19 －

experiment is performed to examine its effect.
Small open ratio of the separation wall resulted
small sloshing height at 1st mode while new modes,
1.5th mode and 2.5th mode, are excited because thetank is divided into 2 parts by the separation wall. 1.5th
mode decrease if we open the center position of theseparation wall and 2nd and 2.5th modes are reduced if
the wall has many small open areas. We find the best
layout of the separation wall from 21 experimental
cases whose normalized sloshing height is restrictedless than 10 in the wide frequency range.
6. Summary
We have conducted research study on precise
evaluation of vulnerability of urban infrastructure
against strong ground motions. Earthquake hazard inJapan is very high and the hypothetical Nankai
subduction earthquakes will occur in near future.
Integration of multiple evaluation indexes of seismic
vulnerability for urban area is continuously important
issue for seismic disaster mitigation. Cooperation
studies of science and technology researchers should
be continued.
Acknowledgements
We used K-NET, KiK-net, and CEORKA strong
motion data.
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－ 21 －