EARTHQUAKE SAFETY EVALUATION OF ATATURK DAM

1 st National Symposium and
Exposition on Dam Safety
May 28-30, 2007
EARTHQUAKE SAFETY EVALUATION OF ATATURK DAM
Martin WIELAND 1 , Sujan MALLA 2
ABSTRACT
The key structure for the development of the Lower Euphrates region is the Ataturk Dam, a
multipurpose dam for hydropower generation, irrigation as well as domestic and industrial water
supply. The 170 m high rockfill dam was designed against earthquakes using design criteria and
methods of analysis that were valid at the time of the design. The construction of the dam was
completed in 1990. A new guideline document on seismic design criteria for large dams was issued by
ICOLD in 1989. As a consequence, the previously used seismic design criteria became obsolete.
Hence, a re-assessment of the earthquake behaviour and safety of the dam was performed for the
maximum credible earthquake (MCE). Based on new detailed information on the seismic hazard in
Turkey, the MCE shaking at the dam site is estimated to have a peak horizontal ground acceleration of
0.5 g. The dam is expected to undergo inelastic deformations during the MCE. The seismic safety
evaluation included the following steps:
• Specification of the ground motion of the MCE.
• Evaluation of the effective static stresses in the dam due to dead and water loads.
• Estimation of strain-dependent material properties for dynamic analysis.
• Dynamic analysis of dam by equivalent linear method using a two-dimensional finite element
model of the maximum section of the dam.
• Analysis of seismic movements of potential sliding masses on the upstream and downstream slopes
of the dam.
• Analysis of permanent inelastic deformations of the dam due to the earthquake shaking.
• Safety evaluation of the dam based on the estimated dam deformations.
Because of the relatively flat slopes, the wide filter zones, and the thick clay core with self-healing
properties, the dam is expected to perform well under the MCE ground motion.
Keywords: Rockfill dam, equivalent linear earthquake analysis, seismic slope stability, dam safety
assessment
INTRODUCTION
The most important results of the earthquake analysis of the Ataturk Dam subjected to the ground
shaking produced by the maximum credible earthquake (MCE) are presented in this paper. The
dynamic analysis of the dam subjected to the earthquake ground shaking was carried out with the help
of a finite element model. Based on the results of this analysis, the dynamic stability and the
earthquake-induced deformations of the dam were examined.
1
Chairman, ICOLD Committee on Seismic Aspects of Dam Design, Poyry Energy Ltd.,
Hardturmstrasse 161, CH-8037 Zurich, Switzerland, e-mail: martin.wieland@poyry.com
2
Senior Engineer, Poyry Energy Ltd., Hardturmstrasse 161, CH-8037 Zurich, Switzerland, e-mail:
sujan.malla@poyry.com

The earthquake analysis of the dam consisted of the following steps:
1. Generation of input ground motion (in the form of artificial spectrum-compatible accelerograms)
corresponding to the MCE, which was selected on the basis of the available information about the
seismic hazard at the dam site
2. Assessment of the following dynamic properties of each dam material
• Maximum dynamic shear modulus (i.e. shear modulus for very small dynamic shear strains),
which depends mainly on the mean effective static stresses and the relative density in the
shells, and on the undrained shear strength in the clay core
• Strain-dependence of dynamic shear modulus and damping ratio
• Shear strength properties: angle of internal friction and cohesion
• Other properties: Poisson’s ratio and mass density
3. Dynamic analysis of the dam subjected to the earthquake ground motion by means of a finite
element model using the equivalent linear method first developed by H.B. Seed and co-workers;
this method is the state-of-the-practice in the earthquake analysis and design of embankment dams
(several cases were analyzed in view of the random nature of the earthquake shaking and
uncertainties concerning the dynamic material properties; the main result of interest from the
dynamic analysis is the distribution of the earthquake accelerations in the dam, in particular, in
potential sliding blocks)
4. Selection of potential sliding surfaces and calculation of yield accelerations by performing slope
stability calculations (note: the yield acceleration of a potential sliding mass is the pseudo-static
horizontal earthquake acceleration for which the factor of safety against a sliding failure is equal
to 1.0)
5. Calculation of the permanent earthquake-induced displacements of the potential sliding masses
based on the knowledge of their yield accelerations and the distribution of the earthquake
accelerations (Newmark sliding block analysis)
6. Seismic settlement analysis to determine the effect of vibration-induced densification of the dam
materials during the earthquake shaking
7. Determination of total loss of freeboard due to the sliding displacements caused by the dynamic
slope instabilities as well as the seismic settlements
8. Assessment of effect of the earthquake-induced deformations on safety against internal erosion in
the dam.
MAIN FEATURES OF ATATURK DAM
The Ataturk Dam is a zoned rockfill dam with a central core and is located on the Euphrates River.
The main features of the dam are as follows:
• Dam height: 170 m
• Crest length: 1670 m
• Crest level: 549 m
• Maximum base width: approx. 900 m
• Dam volume: 84 × 10 6 m 3
• Reservoir volume: 48 km 3
• Installed capacity: 2400 MW
• Annual energy generation: 8100 GWh
The construction of the cofferdam lasted from 1985 to 1987. The fill work for the main dam began in
1987 and was completed in 1990. The reservoir level reached 535 m a.s.l. in March 1994 and has
varied between 526 m and 537 m a.s.l. since then. The maximum and minimum operation reservoir
levels are 542 m and 526 m a.s.l., respectively.

DESIGN EARTHQUAKE GROUND MOTION
The Ataturk Dam is situated about 50 km southeast of the East Anatolian fault. The most recent
damaging seismic activity associated with the East Anatolian fault took place on 27 June 1998 during
the Adana-Ceyhan earthquake, which had a surface wave magnitude of 6.2 and an epicentral distance
of about 270 km from the Ataturk dam site.
To estimate the peak ground acceleration (PGA) of MCE at the dam site, the seismic hazard data from
the Disaster Management Implementation and Research Centre of the Middle East Technical
University (METU/DMC, 1999) were used as a basis. Accordingly, the peak horizontal ground
accelerations at the dam site (rock surface) for various return periods are as follows:
Return period (years) 100 225 475 1000
Peak ground acceleration 0.22 g 0.27 g 0.33 g 0.39 g
The maximum credible earthquake (MCE) is defined as the largest possible earthquake that can be
reasonably expected to occur in the given tectonic setting. The return period of the MCE may be
several thousand years. It is clear that the PGA of the MCE should be larger than that of an earthquake
with a return period of 1000 years. In view of the above-mentioned information about the seismic
hazard at the dam site, the PGA of horizontal shaking at the free surface of the rock due to the MCE
was taken as 0.50 g.
The vertical earthquake acceleration was also considered in this study. The PGA of the vertical
component of the input earthquake motion was assumed to be equal to 2/3 of the PGA of the
horizontal component. The elastic response spectrum defined by in the Turkish code was used
(Ministry of Public Works and Settlement, 1998).
In view of the random nature of the ground motion, three independent sets of artificial accelerograms
were employed in the dynamic analysis. These spectrum-compatible accelerograms were generated
with the computer program SIMQKE (Gasparini and Vanmarcke, 1976). The strong ground motion
(i.e. the stationary part of the input motion) was assumed to last for 25 s, 30 s and 35 s for the three
artificial earthquakes, respectively.
TWO-DIMENSIONAL FINITE ELEMENT MODEL OF HIGHEST DAM SECTION
A two-dimensional finite-element model of the dam was used for the dynamic analysis. Four-node
plane-strain elements were employed. The two-dimensional model represents the central cross-section
of the dam. The model geometry corresponds to the actual shape of the dam surface. The roads built
on the dam surface were also modelled. Some adjustments made in the topmost part of the dam during
the crest reinstatement were also incorporated. The horizontal and vertical degrees-of-freedom were
fixed at the base of the dam, where the input seismic motion defined on the rock surface was applied.
DYNAMIC MATERIAL PROPERTIES
The behaviour of geotechnical materials subjected to an earthquake loading is highly nonlinear. In
particular, the dynamic shear moduli and damping ratios of such materials are strongly influenced by
the amplitudes of the dynamic shear strains. The dynamic characteristics of the dam materials had not
been specially investigated by means of dynamic triaxial tests. Such tests would provide the necessary
information about the dynamic properties of these materials. The material properties required for the
dynamic analysis were obtained from the geotechnical literature on the dynamic characteristics of

comparable materials. Possible ranges of the most important dynamic properties were considered by
analysing a number of cases.
The Poisson’s ratios of the shells (including filters) and the core were assumed to be equal to 0.30 and
0.47, respectively. The total mass densities of the dam materials corresponded to the total unit weights
listed in Table 2.
The principal features of the dynamic properties of the dam materials are explained below.
Maximum dynamic shear modulus in upstream and downstream shells
The maximum dynamic shear modulus Gmax of a cohesionless geotechnical material can be expressed
as (Seed and Idriss, 1970)
0.
5
G max = 220 k 2max ( σ′
m ) (in kPa)
where k2max is a material constant that depends primarily on the relative density and σ′ m is the mean
effective static stress. Conventionally, the shear modulus for a cyclic shear strain amplitude of 10 -4 %
is designated as Gmax, since the dynamic shear modulus is practically constant for strain amplitudes
below this level.
The possible range of values of k2max considered for the upstream and downstream shells was as
follows (Seed et al., 1984):
• Lower bound: k2max = 90
• Average: k2max = 120
• Upper bound: k2max = 150
Maximum dynamic shear modulus in clay core
The maximum dynamic shear modulus Gmax of the clay core depends primarily on the undrained shear
strength su (Seed and Idriss, 1970). The approximate relationship between the two quantities can be
expressed as
⎧1200
lower bound
G max ⎪
= ⎨2400
average
s u ⎪
⎩3600
upper bound
The undrained shear strength su of the clay core varies from approximately 100 kPa at the top to
220 kPa at the bottom of the dam. It is assumed to vary linearly over the dam height.
Strain-dependence of dynamic shear modulus and damping ratio
The dynamic shear modulus and the damping ratio of a fill material subjected to a cyclic loading vary
substantially with the amplitude of the shear strain. In general, the shear modulus diminishes and the
damping ratio grows when the dynamic shear strain amplitude becomes larger. The strain-dependence
of the dam materials for the earthquake analysis is briefly discussed below.
• Upstream and downstream shells: The dynamic shear modulus and the damping ratio of the
upstream and downstream shells (including the filters) were assumed to vary with shear strain
amplitude as shown in Fig. 1. These curves were proposed by Seed et al. (1984) based on a large
number of cyclic triaxial tests made on gravelly soils, including the rockfill materials used in the
Pyramid and Oroville Dams in USA.
• Clay core: The strain-dependence of a clayey soil is strongly influenced by its plasticity index. The
plasticity index of the core material of the Ataturk dam lies within the range of 20 to 45. The
corresponding variation of the shear modulus and the damping ratio of the clay core with the
dynamic shear strain are plotted in Fig. 2, in which the average curves as well as the upper and
lower bounds are indicated.

Fig. 1: Dependence of dynamic shear modulus (left) and damping ratio (right) of rockfill on cyclic
shear strain amplitude
Fig. 2: Dependence of dynamic shear modulus (left) and damping ratio (right) of clay core on cyclic
shear strain amplitude
EARTHQUAKE ANALYSIS
Brief description of analysis method
The dynamic analysis was performed with the computer program QUAD4M (Hudson et al., 1994).
This program evaluates the seismic response of any soil deposit or earth structure using the finite
element procedure in the time domain. The equations of motion are solved by direct step-by-step
numerical integration.
The shear modulus and damping ratio are based on dynamic tests using constant-amplitude strain
cycles. The strain response during an earthquake, however, consists of cycles with variable
amplitudes. This was taken into account by using an equivalent, uniform dynamic shear strain
amplitude equal to 65% of the peak dynamic shear strain for the evaluation of the dynamic shear
modulus and the damping ratio. An iterative procedure, based on the equivalent linear method, was
applied to obtain the strain-compatible shear modulus and damping ratio of each finite element. The
solution was found to converge satisfactorily in 5 iterations.
The integration of the dynamic equations of motion was done in time steps of 0.01 s. The dynamic
solution was obtained for the total duration of the input motion plus an additional 5 s of the free
vibration phase after the earthquake.
Cases analysed

In view of the randomness of the earthquake excitation and the uncertainties concerning the dynamic
soil properties, a number of cases listed in Table 1 were analysed.
Table 1: List of analysed cases and some principal results
Case
Dynamic shear moduli
(Gmax value and G/Gmax
curve)
Damping ratio curve
Fundamental
eigenfrequency
(Hz)
Peak absolute
acceleration at
crest level (g)
Horz. Vert.
A Average Average 0.49 0.64 0.33
B Average Lower bound 0.46 0.75 0.35
C Average Upper bound 0.51 0.55 0.30
D Lower bound Average 0.31 0.42 0.26
E Upper bound Average 0.67 0.93 0.49
Table 1 lists the fundamental eigenfrequencies in the various cases. The fundamental eigenfrequency
of the dam subjected to the MCE lies in the range of 0.31 Hz to 0.67 Hz, and is equal to about 0.50 Hz
for the estimated average material properties. It should be noted that the dam eigenfrequencies also
depend on the earthquake size, i.e. they are smaller for a larger earthquake due to the straindependence
of the dynamic shear moduli.
The peak value of the horizontal component of the absolute crest acceleration is in the range of 0.42 g
to 0.93 g and that of the vertical component in the range of 0.26 g to 0.49 g. The highest and lowest
earthquake accelerations are, respectively, associated with the estimated upper and lower bounds of
the dynamic shear moduli of the embankment materials. This is related to the fact that the spectral
acceleration increases when the fundamental frequency becomes higher in the frequency range up to
3.3 Hz due to the shape of the input response spectrum.
DYNAMIC SLOPE STABILITY ANALYSIS
If the inertial forces acting on a potential sliding mass on a dam slope during an earthquake become
sufficiently large, the total (static plus dynamic) driving forces would exceed the available resisting
forces. As a result, the factor of safety would drop below 1.0, implying that the sliding mass starts to
move because equilibrium is no longer possible. The permanent displacements of such potential
sliding masses on the dam slopes were computed by the Newmark sliding block analysis.
Calculation of yield accelerations
For the Newmark sliding block analysis, first of all, the yield acceleration of a potential sliding mass
has to be determined. This corresponds to the pseudo-static horizontal earthquake acceleration for
which the factor of safety against a sliding failure is equal to 1.0.
The yield acceleration depends mainly on the following factors: (i) gradient of slope, (ii) shape of
sliding surface, (iii) shear strength parameters (angle of sliding friction and cohesion), and
(iv) distribution of pore water pressures.
The slope stability analysis for the calculation of the yield acceleration was performed using the
Spencer method. In this method, the factor of safety is obtained by taking into account both the force
and moment equilibrium conditions, and the resultant interslice forces are assumed to have a constant
slope throughout the sliding mass.

Table 2: Material properties for slope stability calculations
Description
Total unit
weight
(kN/m³)
Cohesion
(kPa)
Angle of friction
Peak Residual
φ't φ'res
Upstream shell (submerged) 24.0 0 43° 39°
Downstream shell and top of upstream shell
above water surface
22.1 0 43° 39°
Upstream filter (submerged) 23.3 0 39° 37°
Downstream filter and top of upstream filter
above water surface
21.9 0 39° 37°
Core (below reservoir level) 20.3 30 10° 10°
Core (above reservoir level) 19.8 30 10° 10°
Note: The φ't values listed above are representative peak friction angles for relatively high confining stresses.
The friction angles for low confining stresses are higher. As the stress-dependence of the friction angles of the
shell and filter materials has not been taken into account, the above friction angles are conservative.
From the viewpoint of dam safety, the sliding surfaces of interest are those involving the dam crest
because any sliding movement in the crest area would lead to a reduction of the freeboard. The
dynamic slope stability calculations were made for a large number of potential sliding blocks on both
the upstream and downstream sides of the dam crest.
The material properties assumed for the slope stability calculations are listed in Table 2. A distinction
is made between the peak and residual values of the angles of friction of the various dam materials.
This reflects the fact the shear strength usually drops to some extent after attaining a maximum value
as shear strains increase.
The maximum operation reservoir level was taken as 542 m a.s.l. for the calculation of pore pressures
in the upstream shell and filter. The pore pressures in the clay core were obtained from the estimated
distribution of pore pressure ratios ru over the dam height (note: ru = u / γ h, where u is the pore water
pressure and γ h is the total vertical pressure expressed as a product of the total unit weight γ and the
depth h).
The yield accelerations of each potential sliding block were calculated for the following three cases:
i) Using peak angles of sliding friction and ignoring influence of vertical acceleration
ii) Using residual angles of sliding friction and ignoring influence of vertical acceleration
iii) Using residual angles of sliding friction and considering influence of vertical acceleration
Newmark sliding block analysis
As soon as the absolute horizontal acceleration of a potential sliding mass exceeds the yield
acceleration, it starts to move. The relative velocity of the sliding mass grows as long as the
earthquake acceleration remains above the yield level. When the acceleration falls below the yield
level, the motion gets braked and, after some time, the sliding mass sticks to the underlying material
again. The permanent displacement due to the earthquake shaking is determined by integrating the
time history of the relative sliding velocity produced by the acceleration pulses exceeding the yield
level. For the Newmark analysis, the time history of the absolute horizontal acceleration averaged over
each potential sliding mass is needed.
The maximum, mean and minimum values of the permanent horizontal displacements of the critical
sliding block at the upstream slope of the dam are listed in Table 3 for the average dynamic material
properties, as well as for the upper and lower bounds of the dynamic material properties. In this table,
the resulting vertical displacements, estimated on the basis of a geometrical consideration, are also

given. For this purpose, the sliding movement was assumed to occur in the direction of the tangent at
the highest point of the sliding surface.
Table 3: Permanent displacements of critical sliding block on upstream slope
(Notation: ayield: yield acceleration; dh: horizontal displacement; dv: vertical displacement)
Peak angles of friction Residual angles of friction
Considering only
Considering only Considering horizontal and
horizontal acceleration horizontal acceleration vertical accelerations
ayield
(g)
Max./
Mean/
Min.
dh
(m)
dv
(m)
ayield
(g)
Max./
Mean/
Min.
(a) With average dynamic material properties
dh
(m)
dv
(m)
ayield
(g)
Max./
Mean/
Min.
Max. 0.41 0.26 Max. 1.03 0.64 Max. 2.34 1.46
0.239 Mean 0.32 0.20 0.185 Mean 0.91 0.57 0.139 Mean 2.03 1.27
Min. 0.19 0.12 Min. 0.79 0.50 Min. 1.66 1.04
(b) Considering possible upper and lower bounds of dynamic material properties
Max. 1.28 0.80 Max. 2.28 1.42 Max. 3.72 2.32
0.239 Mean 0.47 0.29 0.185 Mean 1.07 0.67 0.139 Mean 2.17 1.36
Min. 0.03 0.02 Min. 0.18 0.11 Min. 0.98 0.61
dh
(m)
CONSEQUENCES OF EARTHQUAKE-INDUCED SLIDING DISPLACEMENTS
The most important consequences of the sliding movements induced by the earthquake loading are
discussed below.
Reduction of available freeboard
During a major earthquake, the crest region can be expected to undergo sliding movements towards
both the upstream and downstream sides along a number of criss-crossing failure surfaces. The sliding
movements are likely to be of intermittent nature. Consequently, the displacements towards the
upstream and downstream sides would occur at different times and, thus, can be assumed to have a
cumulative effect on the vertical displacement of the crest region. Hence, the total drop of the level of
the core top was assumed to be equal to the sum of the vertical displacements occurring during the
sliding displacements towards the upstream and downstream sides.
Based on the results for the average properties, the total vertical drop of the core top would be about
1.0 m during the MCE due to the downward vertical displacements of 0.8 m and 0.2 m produced by
sliding towards the upstream and downstream sides, respectively. However, if the dynamic shear
moduli approach the upper bound values (case E), the total lowering of the level of the core top could
become as high as 2.3 m as a result of vertical displacements of about 1.5 m and 0.8 m during the
sliding movements caused by the MCE towards the upstream and downstream sides, respectively.
Besides the permanent sliding displacements resulting from the dynamic slope instabilities, the
earthquake ground shaking also produces a general settlement due to the vibration-induced
densification of the embankment materials. The maximum seismic settlement of the top of the core
due to this effect was estimated to be of the order of 0.5 m.
Considering the effects of both the movements of potential sliding masses and the general seismic
settlement, it was estimated that the lowering of the elevation of the core top could approach the order
of 3 m during the MCE. Therefore, the freeboard with respect to the core top must be sufficient at all
times to accommodate such a reduction without endangering the safety of the dam.
dv
(m)

Danger of leakage through core
Relatively wide filter (transition) zones have been provided on both sides of the core. Except for the
reinstated crest region, the upstream filter has an almost uniform width of about 9 m (measured
horizontally), and the downstream filter has a width varying from about 8 m at the top to more than
15 m at the base. The filter widths are gradually decreased in the reinstated crest region. As a result,
the widths of the upstream and downstream filters reduce to about 6 m and 8 m, respectively, at
elevation 545 m a.s.l.
The wide upstream and downstream filters form a very effective first line of defence against
concentrated earthquake-induced leakage through the dam. Moreover, the core is made of highly
plastic clay (mainly CH), a good core material in which it is unlikely that a great deal of erosion would
occur even in the worst conditions of cracking and leakage.
After the sliding movement caused by the MCE, the filter width at the location of a sliding surface
would still be at least about 6 m. This should be sufficient to ensure that uncontrolled leakage does not
develop through the dam in the critical period following an earthquake.
SUMMARY AND CONCLUSIONS
The main points and the conclusions of this study are as follows:
1. The present study deals with the effects of the earthquake ground shaking only.
2. The dynamic analysis and the slope stability calculations were performed using 2-dimensional
(2-D) models representing the central cross-section of the dam; 3-D effects were not considered.
3. As dynamic test results were not available, the dynamic material characteristics for the earthquake
analysis were selected on the basis of the geotechnical literature. In view of uncertainties
concerning the material properties and the random nature of the ground motion, the possible
scatter of the results was investigated by analyzing a number of cases.
4. On the basis of the currently available information about the seismic hazard at the dam site, the
peak ground accelerations (PGA’s) of the horizontal and vertical components of the maximum
credible earthquake (MCE) were taken as 0.50 g and 0.33 g, respectively. The input motion in the
dynamic analysis consisted of synthetic accelerograms compatible with the elastic response
spectrum of the Turkish earthquake code.
5. The horizontal crest accelerations calculated for the upper and lower bounds of the damping ratios
deviate from those calculated for the average damping ratios by about 15%. For obvious reasons,
the earthquake response becomes smaller when the damping ratios increase.
6. The peak value of the horizontal component of the absolute crest acceleration calculated in the
various cases under the MCE ground excitation is in the range of 0.42 g to 0.93 g and that of the
vertical component in the range of 0.26 g to 0.49 g.
7. From the viewpoint of the earthquake safety of the dam, sliding displacements in the crest region
are of main interest. A large number of potential sliding surfaces involving the crest area were
examined in this study.
8. As the excess pore pressures generated by the earthquake shaking in the upstream shell and the
clay core are likely to be quite small, they have not been considered in this study.
9. The sliding blocks on the upstream slope have significantly lower yield accelerations than those
on the downstream slope, whereby it is assumed that the angle of friction of the submerged
upstream shell is equal to that of the downstream shell.
10. The yield accelerations of the upstream sliding blocks are in the range of 0.20 g to 0.26 g for the
peak friction angles and in the range of 0.16 g to 0.20 g for the residual friction angles, when only
the horizontal earthquake accelerations are considered. If the vertical accelerations are also
conservatively applied in the unfavourable direction, the horizontal yield accelerations of the
upstream sliding blocks could become as low as 0.12 g to 0.16 g.
11. Based on the results obtained using the average dynamic material properties, the total vertical drop
of the core top during the MCE is expected to be about 1.0 m due to downward displacements of
0.8 m and 0.2 m produced by sliding towards the upstream and downstream sides, respectively. In

case the dynamic shear moduli approach the more unfavourable upper bound values, the total
lowering of the core top during the MCE could become as high as 2.3 m as a result of drops of
about 1.5 m and 0.8 m due to sliding movements towards the upstream and downstream sides,
respectively.
12. The ground motion also produces general settlements due to the vibration-induced densification of
the embankment materials, which was estimated as 0.5 m.
13. The total drop of the elevation of the core top caused by the MCE could approach the order of
3 m, when the effects of both the earthquake-induced sliding movements along the slopes and the
general seismic settlements due to the densification of the dam materials are considered. Hence,
the minimum freeboard with respect to the core top should not be less than 3 m. In other words,
the minimum level of the top of the core should be at least 545 m a.s.l., as the maximum operation
reservoir level is 542 m a.s.l.
14. The relatively wide filter zones provide an effective first line of defence against the development
of concentrated leakage through the dam after the earthquake. Moreover, the dam core consists of
highly plastic clay, which is a good core material that is resistant to erosion even in the worst
conditions of cracking and leakage. The minimum filter width at the location of a sliding surface
after the MCE would not be less than about 6 m, which should be adequate to ensure that the filter
zones continue to perform satisfactorily in the critical period following the earthquake.
ACKNOWLEDGMENT
The authors are grateful to DSI for permitting publication of the paper. The investigations and studies
described in this paper were carried out in cooperation with DSI and Dolsar Engineering Ltd. The
contributions of other experts, who have participated in this project and are not listed explicitly, are
greatly acknowledged. The opinions expressed in this paper are those of the authors and are not
necessarily those of DSI.
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