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Site effect on vulnerability of high-rise shear wall buildingsunder near and far field earthquakes

Z.P. Wena, Y.X. Hua, K.T. Chaub,*

a Institute of Geophysics, China Seismological Bureau, Beijing, People’s Republic of China

b Department of Civil and Structural Engineering, The Hong Kong Polytechnic University, Yuk Choi Road, Hung Hom, Kowloon,

Hong Kong, People’s Republic of China

Abstract

Worldwide experience repeatedly shows that damages in structures caused by earthquakes are highly dependent on site condition and

epicentral distance. In this paper, a 21-storey shear wall-structure built in the 1960s in Hong Kong is selected as an example to investigate

these two effects. Under various design earthquake intensities and for various site conditions, the fragility curves or damage probability

matrix of such building is quantified in terms of the ductility factor, which is estimated from the ratio of storey yield shear to the inter-storey

seismic shear. For high-rise buildings, a higher probability of damage is obtained for a softer site condition, and damage is more severe for far

field earthquakes than for near field earthquakes. For earthquake intensity of VIII, the probability of complete collapse ( P ) increases from 1

to 24% for near field earthquakes and from 1 to 41% for far field earthquakes if the building is moved form a rock site to a site consisting a

80 m thick soft clay. For intensity IX, P increases from 6 to 69% for near field earthquake and from 14 to 79% for far field earthquake if the

building is again moved form rock site to soft soil site. Therefore, site effect is very important and not to be neglected. Similar site and

epicentral effects should also be expected for other types of high-rise structures.

q

2002 Elsevier Science Ltd. All rights reserved.Keywords: Seismic vulnerability; Damage probability matrix; Site conditions; Epicentral distance; Distant earthquake; Near earthquake

1. Introduction

It has been repeatedly demonstrated by many strong

earthquakes, including the 1906 San Francisco, the 1957

and 1985 Mexico City, the 1967 Caracas, the 1976

Tangshan, the 1989 Loma Prieta, the 1994 Northridge and1995 Kobe earthquakes, that damages of buildings depend

strongly on the local site response. As early as 1906, during

the great San Francisco earthquake it was realized that

damage was more severe at downtown situated on a softground than the surrounding areas [1]. The 1985 Mexico

City earthquake caused only moderate damages in the

vicinity of the Pacific coast of Mexico, but caused extensive

damages some 350 km away in Mexico City. Structural

damages in Mexico City were also highly selective. Large

parts of the city experienced no damage while areas

underlain by 38– 50 m of soft soil suffered pronounced

damages [2]. During the 1989 Loma Prieta earthquake, the

epicentral intensity is only VIII in the modified Mercalli

scale (MMI), while the intensity is IX in some soft site in

San Francisco, which locates more than 100 km away.

In short, the fact that earthquakes caused extensive damage

in certain areas, and relatively little damage in others,

suggests that local site effects are important [2].

In addition, evidences from many earthquakes repeat-

edly illustrate that damage phenomenon in the near field

and far field are quite different. For example, the 1952Kern earthquake caused heavier damages to one-storey or

two-storey brick buildings than to multi-storey buildings

in the epicentral areas, while caused heavier damages to

five or above multi-storey buildings than to low-risebuildings in Los Angeles about 150 km away [3]. Thus,

the damages are highly selective in terms of both the

natural frequency of structures and the frequency content

of ground shaking.

Although effects of local site condition and epicentral

distance on building damage have been confirmed by many

earthquakes and have been investigated extensively [4–7],

these local site effects have not been incorporated into

vulnerability analysis of buildings through the use of

damage probability matrix (DPM) or fragility curves. To

incorporate the local site effects on the damages of existing

buildings, subjective expert opinions are often being used.

As remarked by Hu et al. [8], Medvedev [9] was perhaps thefirst to summarize systematically the effect of site

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Soil Dynamics and Earthquake Engineering 22 (2002) 1175–1182www.elsevier.com/locate/soildyn

* Corresponding author. Fax: þ852-2334-6389.

E-mail address: [email protected] (K.T. Chau).

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conditions on building damages. On the basis of field data,

including the shear wave velocity and the depth of water

table, he correlated structural damages to the site condition.

For the 1967 Caracas earthquake, Seed et al. [10] correlated

statistically different types of structural damages to the local

soil conditions. Kuribayashi et al. [11] related the

probability of damage of wooden houses during the 1948

Fukui earthquake to the local ground condition. Cochrane

and Schaad [12] presented a simple method to consider the

effect of soil condition on vulnerability of buildings by

either increasing or decreasing the design intensity. By

applying this method, Chavez [13] analyzed the effect of

local geology on the seismic vulnerability of the metropo-

litan zone of Guadalajara, Mexico. More recently, Murao

et al. [14] incorporated the site effects by using fragility

curves or DPM formulated based on the damage survey datafrom the 1995 Kobe earthquake; and Mucciarelli et al. [15]

incorporated the site effects by using microtremor

measurements.

In reality, the local site condition and epicentral distance

may influence the magnitude as well as the frequency

content of strong ground motions. This paper attempts to

analyze the combined effects of the soil condition and

epicentral distance on the vulnerability of a typical

reinforced concrete frame/shear wall building in Hong

Kong by proposing the DPM for various conditions. More

specifically, we will use a multi-degree-of-freedom lump

mass system to represent a 21-storey reinforced concretebuilding in Hong Kong. The shear force will be compared to

the yield shear to estimate the ductility, and subsequently

the damage states of the building.

2. Formulation for damage probability matrix

2.1. Input ground motion

As mentioned in Section 1, the importance of site

condition and epicentral distance on local ground motions is

well recognized. In this study, however, no topography and

basin effect is incorporated. In addition, the effect of

duration of strong ground motion is not included. Instead,

we adopt the site- and earthquake-dependent design

response spectra of the Chinese seismic code GBJ 11-89

[16] as our seismic input. As shown in Fig. 1, the seismic

coefficient a is expressed in spectrum form, and depends on

the natural period of the site T g and the site condition. The

site conditions can be classified into four categories, namely

SC I, SC II, SC III, and SC IV; and they correspond to a stiff site, a medium-stiff site, a medium-soft site and a soft site.

The exact definitions are given in GBJ 11-89. Typically, SC

I is a rock site; SC II corresponds to a site with less than 9 m

thick of stiff soil with shear wave speed vs . 500 m/s; SC III

corresponds to a site with either a 3–80 m thick of medium-

stiff soil with 500 m/s $ vs . 250 m/s or a medium-soft soil

(250 m/s $ vs . 140 m/s) of more than 80 m thick; and SC

IV corresponds to a site with a soft soil with vs # 140 m/s of

more than 80 m thick. The site fundamental period T g can be

estimated from Table 1 as a function of site category as well

as whether the design earthquake is far field or near field,

ranging from 0.2 to 0.86 s. The maximum seismic

coefficient amax given in Fig. 1 depends on the designlevel of earthquake intensity. The values of amax corre-

sponding to MMI VI, VII, VIII, IX and X can be taken as

0.12, 0.23, 0.45, 0.90 and 1.80, respectively.

Fig. 1. Design spectra of GBJ 11-89 for different site conditions at 5% damping. The natural periods of the site and the structure are denoted by T g and T ,

respectively, while the site categories I, II, III, and IV are denoted by SC I, SC II, SC III, and SC IV, respectively.

Z.P. Wen et al. / Soil Dynamics and Earthquake Engineering 22 (2002) 1175–11821176



2.2. Seismic storey shear

By applying the equivalent lateral force method, each

level of a building can be modeled by one lateral degree-of-

freedom along the shaking direction. The lateral force

applied at level i can be calculated from Ref. [16]:

F i ¼ Gi H iXn

j¼1G j H j

aGeqð12 d nÞ ð1Þ

where H i and Gi are the height and weight at level i,

respectively, n is the total number of stories of the building,

a is the spectrum parameter given in Fig. 1, d n is the

additional seismic action coefficient given in Fig. 2, and Geq

is the total equivalent weight of a structure (or 85% of the

total weight of the building). The coefficient d n is introduced

to account approximately for the higher mode contributions,

and such approach is essentially the same as those used in

the UBC-85 of the USA.The shear force Q x at the x storey is then given by

summing all lateral seismic forces above that storey, i.e.

Q x ¼Xn

i¼ x

F i þ aGeqd n ð2Þ

where F i is given in Eq. (1).

In using Figs. 1 and 2, the fundamental period T of a

building is needed, and for RC frame/shear wall buildings T

can be approximated by the following empirical formula

[17]:

T ¼ 0:33 þ 0:00069 H 2 ffiffi B

3p ð3Þ

where T is given in second, H and B are the height and

length along the shaking direction of the building (in m).

2.3. Yield shear coefficient of each storey

For frame structure with shear walls, the yield storey

shear can be estimated as [16]:

Q yx ¼ 0:25F c Awx ð4Þwhere F c is compressive strength of concrete and Awx is

sectional area of shear walls which are parallel to theearthquake action in storey x. Note that this yield storey

shear will not be constant at different levels if the building is

not uniform along the height. This formula is similar to

those adopted in Japan [18]. That is, the yield storey shear

should be independent of the frequency content of the

applied seismic forces, and the ultimate strength is assumed

proportional to the area of the shear wall aligned along the

direction of the seismic loads. The seismic resistant capacity

of a high-rise building can be estimated by a yield shear

coefficient of storey defined as:

R

¼

Q yx

Q x ð5

Þwhere Q x is the seismic shear in the x storey given in Eq. (2)

and Q yx is yield shear of the same storey given in Eq. (4).

Many studies show that, in the case of multi-storey frame

structure with shear walls, non-linear deformation will

concentrate at the weakest stories [19,20], which correspond

to the minimum R in Eq. (5). Note also that in the

calculation of Q x, the dynamic characteristics of the ground

motions have been taken into account, approximately. Thus,

the yield shear coefficient given in Eq. (5) relates not only

to the strength of the structure, but also to the characteristics

of the seismic input.

2.4. Probability density of the ductility factor

The maximum storey ductility factor is a key parameter

indicating building damage. The storey with minimum yield

shear coefficient experiences the maximum deformation and

attains the maximum ductility factor. Based upon 3120

cases of elastic–plastic seismic analyses of different types

of multi-degree-of-freedom structures subject to 31 real

seismic records, Yin et al. [21] proposed the following

formula for the maximum mean ductility m0 of a frame

structure with shear wall:

m0 ¼exp½2:6ð12 RÞ ffiffi

Rp R # 1

1 R

R . 1

8<: ð6Þ

Table 1

Characteristic period of the site T g (s) in terms of site category (SC I, SC II,

SC III, and SC IV) and for both near and far field earthquakes (after GBJ

11-89)

Epicentral distance T g

SC I SC II SC III SC IV

Near-earthquake 0.20 0.30 0.40 0.65

Far-earthquake 0.25 0.40 0.55 0.85

Fig. 2. A plot of the additional seismic action coefficient d n at the top levelof the structure versus the structural period T . The natural period of the site

is denoted by T g.

Z.P. Wen et al. / Soil Dynamics and Earthquake Engineering 22 (2002) 1175–1182 1177



where R is the minimum yield storey shear coefficient

calculated from Eq. (5). This result for ductility factor is

comparable to the formula by Veletson and Newmark [22]

derived from the energy principle [23]. This empirical

formula can further be refined by adding correction factors

C i ði ¼ 1; 2; …; 5Þ to the maximum mean ductility factor:

m ¼ m0 1 þX

C i

ð7Þ

When sectional stiffness is non-uniform along the height,

C 1 ¼ 0.2; when the plan of building is non-symmetric,

C 2 ¼ 0.2; when the quality of the building is sub-standard,

C 3 ¼ 0.2; when the building design complies with require-

ment of TJ11-78 (Chinese seismic code for industrial and

civil buildings, 1978), C 4 ¼ 20.25; and, finally, when

building design complies with TJ11-74 (Chinese seismic

code for industrial and civil buildings, 1974) but not

TJ11-78, C 5 ¼ 20.2. Otherwise, we can set C i ¼ 0

ði ¼ 1; 2; …; 5Þ.Naturally, a higher maximum mean ductility factor m

implies a more severely damaged building. But, due to the

uncertainties involved in the estimation of seismic hazard

as well as in the analysis of structural response, the ductility

as well as the damage state is better represented in terms of

probability distribution function. In particular, both the

peak ground acceleration and seismic capacity of structures

are often found satisfying a lognormal probability distri-

bution versus the input ground parameters [14,24–27].

Indeed, by analyzing 3120 cases of elastic–plastic seismicresponses, Yin [24] suggested the following lognormal

distribution for m

f ðmÞ ¼ 1 ffiffiffiffi2p

p jm

exp 2ðln m2 lÞ2

2j 2

" #ð8Þ

where

l ¼ ln m2 12j 2; j 2 ¼ ln 1 þ s 2

m2

!ð9Þ

In these equations, m and s are, respectively, the maximum

mean value estimated from Eq. (7) and the standard

deviation of the ductility factor of the stories. In this study,we assume that the main uncertainty of the ductility factor

comes from uncertainty in the seismic hazard. Thus, thevalue of s = m is taken from that of the earthquake intensity

as 1.25 [28]. Note that this ductility distribution depends on

the structural characteristics as well as the input ground

motions through the calculation of m:

2.5. Damage probability matrix

Fig. 3 shows a typical plot of the base shear Q versus the

ductility factor. Five damage states, D1, D2, D3, D4 and D5,

are assumed, corresponding to undamaged, slightly

damaged, moderately damaged, extensively damaged, andcompletely damaged states, respectively. For frame

structure with shear walls, the threshold ductility factorsfor the onset of slightly damaged, moderately damaged,

extensively damaged and completely damaged states are

1.0, 1.5, 3.0 and 5, respectively [24]. These values are

empirical constants and may vary from one type of building

to another.

Because of the randomness of ground motions and

seismic responses, the earthquake damages are modeled asprobabilistic phenomena in order to closely reflect its

scattering nature of occurrence. Damage probability distri-

bution for various damage states of a specific building isrepresented as the DPM. It describes the probability that the

structure is in a particular damage state for a given level of

ground shaking. Using the threshold values of the ductility

factor as integration limits, the probability of various

damage states for a given ground motion can be obtained

by integrating Eq. (8).

In Section 3, a particular structure will be used as an

example to illustrate the combined effect of the soil

condition and epicentral distance on DPM for various

design intensities.

3. DPM for a 21-storey RC frame/shear wall building

One particular 21-storey RC frame/shear wall building of

Mei Foo Sun Chuen is chosen for our study because Mei

Foo Sun Chuen is one of the largest private housing estates

in Hong Kong. More importantly, the structural scheming

for Mei Foo Sun Chuen is also similar to those used in other

newer residential buildings built recently in Hong Kong.

There are a total of 102 building blocks in Mei Foo Sun

Chuen, and all of them are of the same height and of similar

structural scheming. The whole estate was built in eightdifferent phases on a reclaimed land in the 1960s and 1970s.

Fig. 3. A schematic diagram illustrating a typical plot of the base shear Q

against the ductility factor m. The segments of the curve representing

various damage states D1, D2, D3, D4 and D5 (corresponding to nodamage, slight damage, moderate damage, extensive damage and complete

damage) are indicated. The thresholds of the ductility factor corresponding

to the onset of slight damage, moderate damage, extensive damage and

complete damage are 1.0, 1.5, 3 and 5, respectively.




The building that we choose for the present study is the 5-7

Humbert Street of Phase 5, as shown in Fig. 4(a). The lower

two levels are car parks and the upper 19 levels are

residential stories, and these two levels are connected at the

podium level. There is a sudden change in the building

stiffness from the upper to the lower levels. The structural

plan of the columns and shear walls for a typical storey of

the upper level is shown in Fig. 4(b), while the two car park

levels shown in Fig. 4(c). The storey height and area of

storey at various levels are compiled in Table 2. The height

of the building H is taken as 55.75 m while the length is

taken as 32 m, and the corresponding natural period of the

structure estimated from Eq. (3) is about 1 s.

For our parametric studies, the building is assumed to

rest on different site conditions SC I, SC II, SC III, and SC

IV, and subject to various levels of seismic intensities (VI,VII, VIII, IX and X of a return period of 475 years). Both far

field and near field earthquake excitations are considered.

The self-weight of the building at various level is calculated

based on a distributed load of 1.2 ton/m2. Although the

structure is neither perfectly symmetric nor perfectly

uniform along the height, its stiffness variation along the

height is considered not too drastic. Also noting the fact that

the building is about 30 years old, thus a combined total

correction of 0.25 is applied toP

C i in Eq. (7).

Applying formulas (1)–(7), the maximum mean ductility

m for the building subject to various earthquake intensities,

site conditions and near or far field earthquake is tabulated

in Table 3. Table 2 also complies the storey yield shear atvarious levels calculated according to Eq. (4). As expected,

m increases with the intensity. In addition, m is larger for far

field earthquakes than for near field earthquakes, and this

agrees with the field observation that high-rise building is

more responsive to long period far field ground motion. The

integration of different segment of ductility curves leads to

the probability of various damage states (Fig. 3). For various

site conditions, Table 4 tabulates the DPM of the selected

building for the five damage states (D1, D2, D3, D4 and D5)

subject to intensity levels from VI to X under both far field

and near field earthquakes. Table 4 demonstrates that the

distribution of the ductility factor and the DPM depend

strongly on the site condition and on whether the earthquake

is near or far field.

For the selected building, Table 4 shows that a higher

probability of damage is obtained for a softer site condition.

For intensities VI and VII, the site condition is not very

significant for both near and far field earthquakes. For

earthquake intensity of VIII, the probability of complete

collapse (P ) increases from 1 to 24% for near field

earthquake and from 1 to 41% for far field earthquake if the building is moved form a rock site (SC I) to a site of

80 m thick of soft clay (SC IV). For intensity IX, P increases

from 6 to 69% for near field earthquake and from 14 to 79%

for far field earthquake if the building is moved form SC I

site to SC IV site. For typical reclamation sites in Hong

Kong consisting of 40 m of fill, alluvium, and marine

deposit, the site condition can be approximated by SC II;

and thus, the probabilities of structures suffering from

moderate damage to complete collapse are 1, 18, 70 and

96% for near field earthquake, and 9, 26, 85 and 98% for far

field earthquake for earthquake intensity of VII, VIII, IX, X,

respectively. Therefore, site effect is very important and notto be neglected. In short, a high-rise RC building resting on

a soft site is more conducive to damages than to rest on a

rock site, as it has been demonstrated in the case of 1985

Mexico City earthquake. In addition, high-rise building is

more conducive to far field earthquake than to near field

earthquake. It is because far field seismic ground motions

are richer in higher period content than the near field seismic

motions. This conclusion also agrees with the field

observations during large earthquakes.

Another popular way to present the DPM is to plot the

exceeding probability for a particular damage state (i.e.

probability of a specified damage level will be exceeded)

versus the input seismic intensity as vulnerability or fragilitycurves, as used in Shinozuka et al. [29] and Karim and

Fig. 4. (a) A photograph of the selected 21-storey RC frame-shear wall

building in Mei Foo Sun Chuen used in the vulnerability analysis; (b) theplan section of the columns and shear walls of the building shown in (a);

and (c) an enlargement of the first two levels of parking.

Table 2

The storey height, area and yield shear for the selected frame-shear wall

building from Mei Foo Sun Chuen

Storey no. Storey height (m) Area of storey (m2) Qyx (MN)

20 2.65 548.1 5.08

11 –19 2.65 548.1 18.9

10 2.54 548.1 18.9

9 2.77 548.1 24.3

6– 8 2.65 548.1 24.3

4– 5 2.65 548.1 29.2

3 2.74 548.1 29.2

Podium 2.71 882.2 27.4

Upper parking 2.59 940.4 35.3

Ground parking 2.59 907.1 38.1




Yamazaki [30] for bridges. But such fragility curves will not

be given here as they can be generated readily from Table 4.

Although only one particular high-rise building has been

employed as an example in this study, we expect that similar

conclusions can also be obtained of other high-rise buildings

in Hong Kong. We also expect that the results of this paper

should not be very sensitive to the particular base shear

method or seismic code being adopted.

The results in this paper indicate that damages of the high-

rise buildings are more severe in soft site under far field

earthquakes than in stiff site under near field earthquakes.

This is consistent with commonly observed phenomenon

of ‘highly selective-damage’. Tall flexible structures (with

a long natural period) on soft soil are more conductive to

damage under far field large earthquakes than on rock site

under near field earthquakes. On the contrary, stiff low-rise

buildings appear to be vulnerable to earthquake when they sit

on firm soil under near field moderate earthquakes. There-

fore, the damages are highly selective in terms of the natural

frequencies of structures and of ground shaking.

Physically, large earthquake is more capable of produ-

cing longer-period ground motions than smaller earthquake

does. As seismic waves travel along the Earth crust from a

fault, their higher-frequency components are scattered and

Table 3

The maximum mean storey ductility factor versus the design intensity (VI, VII, VIII, IX, and X) for various site conditions (SC I, SC II, SC III, and SC IV) and

for both near and far field earthquakes

Intensity The maximum mean ductility factor of storey ð mÞ

SC I SC II SC III SC IV

Near Far Near Far Near Far Near Far

VI 0.185 0.226 0.263 0.34 0.34 0.445 0.519 0.663

VII 0.354 0.433 0.5 0.65 0.65 0.855 0.994 1.1

VIII 0.688 0.848 0.978 1.306 1.306 2.79 3.95 6.375

IX 1.7 2.875 4.00 6.65 6.65 10.45 12.86 17.36

X 7.75 10.6 13.0 17.9 17.9 24.05 34.88 34.16

Table 4

DPM in probability of damage (P ) versus intensity for site conditions SC I, SC II, SC III and SC IV (Near: near field earthquakes; Far: far field earthquakes)

Damage state Probability of damage (%) by intensity

VI VII VIII IX X

Near Far Near Far Near Far Near Far Near Far

SC I

D1 98 98 94 91 81 75 48 27 5 3

D2 2 2 4 5 9 11 16 16 6 3

D3 0 0 2 3 8 10 22 26 20 15

D4 0 0 0 1 1 3 8 17 20 18

D5 0 0 0 0 1 1 6 14 49 61

SC II

D1 97 95 95 82 70 58 18 7 2 0

D2 2 3 4 9 12 16 12 8 2 2D3 1 2 1 7 13 17 28 21 12 7

D4 0 0 0 2 4 6 18 22 16 11

D5 0 0 0 0 1 3 24 42 68 80

SC III

D1 95 90 82 74 58 28 7 3 0 0

D2 3 6 9 12 16 16 8 5 2 1

D3 2 3 7 10 17 27 21 13 7 4

D4 0 1 2 3 6 15 22 19 11 8

D5 0 0 0 1 3 14 42 60 80 87

SC IV

D1 87 81 68 65 18 8 2 1 0 0

D2 7 9 14 14 12 8 2 1 1 0

D3 5 8 13 15 28 23 12 7 2 2

D4 1 2 3 4 18 20 15 12 4 5D5 0 0 2 2 24 41 69 79 93 93




dissipated more rapidly than their lower-frequency com-

ponents (see Ref. [31] or fig. 3.23 of Ref. [2]). As a result,

the frequency content also changes with the epicentral

distance, and, thus, far field earthquakes are richer in longer-

period motions. Therefore, tall flexible buildings with

longer natural period are more vibrant when subject to far

field earthquakes.

4. Conclusion

In this paper, the site and epicentral distance effects on

the vulnerability of high-rise buildings were incorporated

into the DPM. This is the first time that these effects are

incorporated into the DPM following a systematic and

analytical approach, instead of using a more subjective‘expert adjustment approach’. More specifically, we have

used a multi-degree-of-freedom lump mass system to

represent a typical high-rise reinforced concrete building

in Hong Kong. The seismic forces at each storey level are

calculated using the base shear method. The effects of the

site condition and the natural period of the building are

automatically incorporated by using a site- and earthquake-

source-dependent design response spectrum. In addition, the

chosen seismic input parameters also depend on whether the

design earthquake is near field or far field. This seismic

force is then compared to the yield shear of each storey,

such that a storey yield shear coefficient can be obtained.

This coefficient is subsequently used to estimate themaximum mean ductility of the building through the use

of a simple empirical relation. The adjusted maximum mean

ductility is then used to form a lognormal distribution of the

ductility factor. By integrating these distributions, fragility

curves or DPM of the building under various types and

levels of seismic input have been obtained. The results in

Table 4 show that damage for high-rise buildings is more

severe for far field earthquake than for near field earthquake

because of its richer low frequency contents. Therefore,

high-rise buildings on soft soil and subject to far field

earthquakes are more conducive damages than on rock site

and subject to near field earthquakes. This is consistent with

commonly observed field phenomenon of highly selective-damage. We expect that the same conclusion can also be

drawn even when we use another high-rise building and

follow a slightly different analytical method.

In conclusion, both the effects of site condition and

epicentral distance should be incorporated analytically into

the calculation of DPM or fragility curves. Site and

epicentral distance effects are important and not to be

neglected.

Acknowledgements

The research was supported by ASD projects A202and A214 of the Hong Kong Polytechnic University.

The authors would like to thank Mr Philip Kwok the

Buildings Department of Hong Kong SAR Government in

providing building information.

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