cfrd 05 bmaterón

RD 2011 CFRD 05

CONSIDERATIONS ON THE SEISMIC DESIGN OF HIGH CONCRETE

FACE ROCKFILL DAMS (CFRDs)

Bayardo Materón

1,, Gabriel Fernandez

2

1, Director Bayardo Materón Associados Ltda, Av. Giovanni Gronchi, 5445- s/172, CEP- 05724-003, S.Paulo -

Brazil,

e-mail: [email protected]

2, Independent Consultant, 207 Sherwin Circle Urbana, IL 61802 – USA

e-mail: [email protected]

ABSTRACT

The height of Concrete Face Rockfill Dams (CFRDs) has been steadily increasing during the past forty years.

During this period the height of CFRDs have increased from 110m (Cethana, Australia, 1971) to 233m

(Shuibuya, China, 2008) and higher dams are under study. CFRD design has gradually evolved from a

predominately empirical approach based on precedent, as well as field observations and measurements, to the

recent use of numerical modeling. However, analytical models to discern static and dynamic behavior have been

absent. This paper focuses on the seismic design of CFRDs taking into account documented field observations

and basic principles of seismic response. The paper summarizes the observed behavior of two dams under

significant seismic loading, the Zipingpu CFRD of compacted rockfill in China shaken by the Wenchuan Quake

(5/12/2008) and the Ishibuchi CFRD of dumped rockfill in Japan affected by the Iwate Miyasi Quake

(6/14/2008); and discusses some aspects of CFRD dynamic response including a simplified approach to evaluate

seismic – induced displacements based on the concept originally proposed by N. Newmark (1965). Suggestions

are also provided on design adjustments to minimize deleterious effects on embankment behavior subject to

large seismic loads that the authors are already applying in high CFRD structures under design and construction.

Key words: CFRDs, Seismicity, Dynamic Response, Design.

1. INTRODUCTION

The stability of embankment slopes subjected to seismic loads was initially treated as a pseudo-static

problem by estimating a “dynamic factor of safety” which included a dynamic force that was

approximated as a fraction (5% to 20%) of the weight of the sliding mass. The meaning of the

dynamic factor of safety however was somewhat nebulous, because the magnitude of the dynamic

forces is not constant and their direction changes throughout the duration of the quake. Furthermore, a

dynamic factor of safety of 1.0 would not preclude the development of seismic induced displacements.

An estimate of the seismic-induced slope displacements after shaking is required for the engineer to

make a judgment on the potential for damage to the dam, the likelihood for a decrease in the shear

strength of the embankment materials or an unacceptable loss of freeboard. A simplified approach to

evaluate seismic-induced displacement in embankment dam was developed by N. Newmark (1965). A

procedure to evaluate seismic response of embankment dams was subsequently developed by

Ambraseys and Sarma (1967) to estimate ground motions amplifications throughout the embankment

taking into account its fundamental period of vibration and the damping of the embankment materials.

Afterwards, a method for estimating embankment earthquake-induced deformations was proposed by

Makdisi and Seed (1978), incorporating the seismic response of the embankment and using the

concept originally proposed by N. Newmark (1965) to estimate seismic-induced displacements.

Numerical modeling is extensively used in current engineering practice to assess embankment

response during seismic loading. This methodology is a useful resource which provides for a

comprehensive evaluation of embankment performance. However, it is always prudent to have

The Second International Symposium on Rockfill Dams

2

analytical models (from shaking experiences) to calibrate and confirm the validity of the numerical

model results. Analytical models can also allow a readily assessment of key parameters controlling

embankment behavior. This paper describes a simplified procedure to assess the seismic response of

rockfill embankments using the concepts originally proposed by Ambraseys and Sarma and N.

Newmark, but with some modifications applicable to CFRD’s. An objective of this paper is also to use

this simplified model to evaluate the beneficial effects of design adjustments currently being

implemented in the design of high CFRDs in seismically active zones and to explore the

implementation of additional design measures.

2. PRECEDENT OF CFRD SEISMIC BEHAVIOR

Prior to 2008 the experience of CFRDs behavior under heavy seismic loading was relatively limited

although the performance of central core rockfill dams under seismic loads had been extensively

documented (Marsal and Ramirez, 1967). One of the early documented cases of CFRD performance

under seismic loading, Cogoti Dam in Chile, was presented by Arrau, et al (1985). The dam is a

dumped rockfill concrete face dam, 85 m high, which was subjected to a PGA of 0.19 g triggered by

the Illapel earthquake on April 4th, 1943, with a magnitude of 7.9, and an epicenter at approximately

90 km from the dam. The crest of the dam is 8m wide and the slopes are 1.4 H:1V and 1.5 H:1V for

upstream and downstream respectively. The dam crest settled 40 cm or 0.47% of its height. The face

slab was designed and built with panels of 10 m x 10 m and with a variable thickness from 20 cm at

the crest to 80 cm at the bottom of the dam. Vertical joints had a 1 inch thick bituminous filler. No

major damage was observed in the face slab, although some deterioration of the joints and spalling of

the central joints have been observed during the life of the project.

In 2008 two major earthquakes occurred in China and Japan which were close to significant CFRD

structures. The China Wenchuan earthquake on May 12, 2008 with a magnitude of 8 had an epicenter

located at 17 km from the Zipingpu CFRD, 156 m high, resulting in a PGA of about 0.5 to 0.6 g. The

dam performance has been discussed by several authors including Chen (1990), Xu (2009) and

Wieland (2010). The crest of the dam is 12 m wide and the upstream slope is 1.4H:1V. The

downstream slope was designed with two inclinations, changing from 1.4H:1V at El.840, to 1.5H:1V

to the crest of the dam at El. 884m as a provision for an eventual earthquake shaking. Fig 1 shows a

cross section of the dam with the zoning following the international nomenclature. The crest settled 74

cm or 0.47% of its height and the maximum downstream displacement was reported to be close to 20

cm, at the crest of the dam.

Fig. 1 Cross Section of Zipingpu CFRD showing the International Nomenclature.

12,0m EL. 884m

EL. 840m

EL. 877m

EL. 830m

2B3A Filter 3m thick

Concrete facing slab

Dumped fill

EL. 728m

Curtain grouting

Original ground surface

3B Rockfill

3C Rockfill3D Rockfill

1.51.0

1.41.0

1.41.0

EL. 763m

Foundation

Interbedded sedimentary rocks

1.41.0

Zipingpu Dam


At the time of the shaking the reservoir was low at El. 830 m about 47 m below the full supply level at

El. 877 m, and only 68% of the water head was app

the upstream slope was confined by the reservoir head while the upper zone of the slope (from El. 884

m to El. 830 m) was unconfined during the severe shaking. Recorded peak accelerations at the crest

the dam near the center of the valley are presented by Ishihara (2010) as follows:

Perpendicular to axis

2.061 g 1.635 g 2.065

Figures 2 to 5 show photographs of the slab above the reservoir level as well as of the downstream

slope. The construction joint between stage 2 and stage 3 at El. 845 m (Fig. 2) was seriously damaged.

It was also observed that the perimeter joint

affected at the central part of the dam.

Fig. 2 Horizontal Joint Damaged at El.845 Fig. 3 Perimetric

Some spalling is reported at the left abutment and at the center of the dam. The severe shaking

triggered cracking at the crest as seen in Fig. 4 and loosening of rock boulders in the downstream

slope Fig. 5, but the performance of the dam from the view point of stability was adequate. The

Zipingpu dam was built in 2006, following the state of art of CFRD’s in China. The dam had an

adequate performance under an earthquake loading higher than the original design, indicating th

resilience of well designed CFRDs under dynamic loading and emphasizing the importance of

including safety measures for the future higher dams.

Fig.4 Some Cracks at the Crest Fig. 5 Rockfill Loosened at

The Miyagi Inland earthquake with a magnitude of 7.2 had the epicenter located at 10 km of the

Ishibuchi CFRD, a dumped rockfill dam, 53 m high, as described by Matsumoto et al (2010),

Yamaguchi (2008) and Wieland (2010). The crest of th

was steep, varying between 1.4 H:1V to 1.2 H:1V at the upper sections of the dam. The downstream

slope was constant with two berms and slope of 1.5 H:1V as shown on Fig 6.

3 stage

2 nd stage

El.845



El. 877 m, and only 68% of the water head was applied on the face slab (Fig 1). Thus the lower part of


m to El. 830 m) was unconfined during the severe shaking. Recorded peak accelerations at the crest

the dam near the center of the valley are presented by Ishihara (2010) as follows:

Parallel to axis vertical up and down

2.061 g 1.635 g 2.065 g



It was also observed that the perimeter joint (Fig. 3), between the face slab and parapet was intensively

affected at the central part of the dam.

Fig. 2 Horizontal Joint Damaged at El.845 Fig. 3 Perimetric Joint between Slab and Parapet



ance of the dam from the view point of stability was adequate. The


adequate performance under an earthquake loading higher than the original design, indicating th


including safety measures for the future higher dams.

Fig.4 Some Cracks at the Crest Fig. 5 Rockfill Loosened at



Yamaguchi (2008) and Wieland (2010). The crest of the dam is only 6 m wide and the upstream slope


slope was constant with two berms and slope of 1.5 H:1V as shown on Fig 6.

2 nd stage slab

3


lied on the face slab (Fig 1). Thus the lower part of


m to El. 830 m) was unconfined during the severe shaking. Recorded peak accelerations at the crest of

vertical up and down



(Fig. 3), between the face slab and parapet was intensively

Joint between Slab and Parapet



ance of the dam from the view point of stability was adequate. The


adequate performance under an earthquake loading higher than the original design, indicating the


Fig.4 Some Cracks at the Crest Fig. 5 Rockfill Loosened at Downstream Slope



e dam is only 6 m wide and the upstream slope


parapet


4

Fig. 6 Cross Section of Ishibuchi CFRD

The average seismic-induced settlement of the crest was 60 cm or 1.13% and the corresponding

maximum downstream displacement was not clearly reported but it was indicated that the downstream

deformations were much higher than the upstream. During the earthquake the reservoir was also low

at El. 300, eighteen meters under the full supply level at El. 318, or 63% of the maximum water head

applied on the face slab.

During construction of the dam, 9 pillars of reinforced concrete were built to provide a bridge for

dumping rockfill. These pillars remained in the rockfill when the dam was completed, and the

settlement of the crest was more prominent between pillars reaching values up to 80 cm settlement or

1.5% deformation, almost 3 times larger than at the Zipingpu dam in China.

Fig. 7 Construction of Ishibuchi Dam by Dumping Rockfill from a Bridge Supported by Pillars

The peak acceleration perpendicular to the axis was reported as 0.95g at the crest of the dam. Several

boulders near the crest, on the downstream slope fell down. No damage was reported for the face slab

but cracks and distresses appeared on the crest as Zipingpu. The dam was not subject to problems of

stability. Figures 8 and 9 show distress in the dam crest.

Fig. 8 Longitudinal Crack on the Crest Fig. 9 Distresses on the Crest

68,53m 83,53m

6,00m

1 : 1,2

1 : 1,3

1 : 1,4

6,0mEL. 323m

EL. 313m

EL. 299m

EL. 318m

EL. 300m

EL. 270m

Grout Curtain

1.51.0

1.51.0

1.51.0

1,40

2,90

Reinforcedconcrete slab

Base Rock Drainage conduit

Dumpedrockfill

Ishibuchi dam

Min. Water level

Max. Water level


5

3.0 SEISMIC RESPONSE OF CFRD.

Concrete Face Rockfill Dams (CFRDs) are inherently well suited to accommodate significant seismic

loading because:

· Embankment materials are generally sound, well compacted, dense rockfill with high friction angles

and adequate modulus of compressibility. In addition, the relatively tight concrete face, combined

with well executed grout curtains and proper embankment zoning maintains the water level at very

low elevations within the embankment materials, and thus pore pressures within the embankment

are nil; and

· Deformations observed in well compacted rockfill subject to heavy shaking are nominal, lower than

one meter.

However, as CFRD’s become higher, some aspects of their dynamic response deserve special attention

including the larger amplification of ground motions which might trigger the development of

significant seismic-induced displacements in the upper dam sections.

3.1 Ground Motions Amplification.

CFRD’s do not behave as rigid bodies when excited into oscillation by strong earthquakes. Their

response depends on the nature of the ground movements, the properties of the rockfill materials, and

the geometry of the embankment. The combination of all these variables can result in accelerations

within the dam substantially larger than the maximum acceleration at the ground surface. Seismic

performance analyses, Ambraseys and Sarma (1967) as well as actual field observations indicate that

for a given dam, ground motion amplifications increase from the base to the crest of the dam; with

crest stations recording accelerations two to three times larger than those reported at the base of the

dam (Marsal & Ramirez, 1967) and (Resendiz and Romo, 1982). A simplified method to approximate

amplification factors at any location within the embankment was proposed by Ambraseys & Sarma

(1967), assuming that the structure would deform in simple shear, in one dimension. Under these

conditions, the fundamental period of the dam, To, can be approximated as:

To = 2.61 �

��

Where h is the height of the dam and Vs is the shear wave propagation velocity at strain levels

compatible with those induced by the ground shaking on the embankment materials. The Vs value can

be extrapolated from shear wave velocity measurements in the embankment materials. In our

experience, well compacted, dense rockfill materials with unit weights γ ≈ 2.2 T/m3 have Vs values in

the range of 1500 ft/sec (457m/sec) to 2000 ft/sec (610m/sec). The acceleration An at any point within

the embankment can be estimated as: An = Ag kn, where Ag is the maximum ground acceleration;

and kn is the amplification coefficient, which is a function of the fundamental period of vibration of

the dam, To, the degree of damping within the embankment materials, and the height above the base of

the dam. Relationships between the amplification coefficient, kn, and the period of vibration of the

dam, To, at various locations above the base of the dam, assuming a damping coefficient of 20% for

the embankment materials, are shown in Figure 10 from Ambraseys & Sarma (1967). As indicated in

Figure 10, the location within the dam at which the kn is estimated is specified by a single parameter

n = y/h where y is the depth of the point below the crest and h is the height of the embankment. For

embankment materials with damping coefficients other than 20% the amplification coefficients

obtained from Figure 10, need to be multiplied by a correction coefficient β, which can be obtained

from Figure 11 (Ambraseys & Sarma, 1967).


6

Fig.10 Maximum Simultaneous Seismic Coefficient for 20% Damping

Fig. 11 Damping Correction Factor Corresponding to Figure 10

(From Ambraseys and Sarma, 1967)

For example, for a CFRD 600 ft high, with dense rockfill materials with a Vs of 1500 ft/sec to 2000

ft/sec the fundamental period, To, ranges from 0.8 to 1 seconds. The amplification coefficient kn of the

ground acceleration at a point located 120 ft below the crest (n = 0.2) can be estimated as 2.0 for a

20% damping coefficient. For a more realistic damping coefficient of 10%, this amplification factor

increases to 2.6. The kn versus To relationships in Figure 10 indicates that seismic-induced

displacements are more likely to take place along relatively shallow rock wedges, near the crest, where

ground motion amplification is likely to be higher.

3.2 Evaluation of the Dynamic Resistance.

Upstream slip surfaces are not likely to develop in CFRD at normal reservoir elevations, due to the

large resistance force generated by the hydraulic pressures against the tight concrete face. The factor

of safety against upstream sliding is generally an order of magnitude larger than that of equivalent

downstream slip surfaces and thus, the potential for seismically-induced upstream displacements is

concentrated in a shallow zone above normal pool elevation. Seismic-induced displacements however


can occur anywhere on the downstream slope, where the resistance to seismic

on the inclination of the slope and the friction angle of the embankment materials.

The first step in the approach proposed in this analysis involves the calculation of the factor of safety

of the slope under static conditions. For the rockfill embankment in Fig. 12 the static factor of safety

against sliding along a shallow slip surface, AC making an angle

estimated using the analogy of the rigid body resting on an inclined surface as shown in Fig. 13. The

static equilibrium force polygon of the rigid body with a weight W in Fig. 13 equal to the weight of

the wedge ABC, resting on a slip surface with an inclination

Where N = W cos α1; replacing terms: FS =

and where α1, the inclination of the slip surface is equal to the mobilized friction angle,

to maintain static equilibrium, and

Fig. 12 Potential Sliding Wedge Geometry

The magnitude of the static factor of safety of the potential sliding wedge provides an index of the

magnitude of the seismic-induced displacements. In areas of high seismicity it is recommended to

provide a static factor of safety of at least 1.5 to maintain seismic

acceptable values. Sliding wedges with static factors of safety ranging between 1.5 and 1.3 are likely

to experience considerable seismic

ground shaking; and sliding wedges with factors of safety lower than 1.3 are likely to experience

unacceptable displacement under dynamic loading.

After the static stability of the embankment has been determined, th

induced displacements of the embankment slope is to evaluate the dynamic resistance of the slope,

along potential sliding surfaces to identify the critical slip surface which yields the largest dynamic

displacements. The dynamic resistance is defined as the minimum seismic

trigger displacements along the slip surface, and is defined as a vector NW where N is a coefficient

and W is the weight of the potential sliding mass. The direction of the dynam

determined by the minimum reserve shearing resistance in addition to that required for static stability,

which can be mobilized to resist the effect of a dynamic load. Once the dynamic resistance is

determined it is possible to cal

acceleration in the direction of NW which will just overcome the reserve resistance of the sliding

mass.

In the third step, the maximum acceleration, A

the potential sliding mass can be compared with the corresponding “yield acceleration” Ng required to


can occur anywhere on the downstream slope, where the resistance to seismic-induced loads depends


approach proposed in this analysis involves the calculation of the factor of safety


against sliding along a shallow slip surface, AC making an angle α1 with the horizontal, can be



on a slip surface with an inclination α1 can be established as:

FS = � �� ϕ

�� α�

; replacing terms: FS = �� ϕ

�� α�

, the inclination of the slip surface is equal to the mobilized friction angle,

to maintain static equilibrium, and is the friction angle of the rockfill materials.

Fig. 12 Potential Sliding Wedge Geometry Fig. 13 Force Polygon of Sliding Wedge


induced displacements. In areas of high seismicity it is recommended to

e a static factor of safety of at least 1.5 to maintain seismic-induced displacements within


to experience considerable seismic-induced displacements that might require extensive repair after


unacceptable displacement under dynamic loading.

After the static stability of the embankment has been determined, the next step in estimating seismic



dynamic resistance is defined as the minimum seismic-induced force that will


and W is the weight of the potential sliding mass. The direction of the dynamic resistance vector is



determined it is possible to calculate the “yield acceleration” Ng, which represents the threshold


In the third step, the maximum acceleration, An, induced by ground shaking on the center of gravity of


7

induced loads depends


approach proposed in this analysis involves the calculation of the factor of safety


with the horizontal, can be



can be established as:

, the inclination of the slip surface is equal to the mobilized friction angle, m, required

is the friction angle of the rockfill materials.

Fig. 13 Force Polygon of Sliding Wedge


induced displacements. In areas of high seismicity it is recommended to

induced displacements within


at might require extensive repair after


e next step in estimating seismic-



induced force that will


ic resistance vector is



culate the “yield acceleration” Ng, which represents the threshold


, induced by ground shaking on the center of gravity of



overcome the reserve resistance along the assumed sliding surface. The magnitude of the maximum

earthquake-induced acceleration, A

(Ambraseys and Sarma, 1967), and described above. If the maximum acceleration A

sliding mass is lower than Ng, then the slope is safe an

place. On the other hand, if the induced acceleration, A

exceeds the value of Ng, relative displacements will take place along the assumed failure surface. The

magnitude of the accumulated displacements that take place along the sliding surface during these

periods of time when the seismic

using the simplified method proposed by N. Newmark (1965), des

The magnitude and direction of the Dynamic Resistance Force NW can be determined using the force

polygon in Fig. 14 which shows the forces acting on a rigid block, resting on an inclined base

representative of the potential sliding mass in

and direction of the weight vector W are known. The mobilized friction angle,

maintain static equilibrium along the sliding surface is also known as and is equal to

of the assumed sliding surface. The direction of the resultant force, R, on the sliding surface is also

known, and is determined by the friction angle,

The friction angle, , of well compacted rockfill materials, generally ranges between 45 to 50

degrees, and can be approximated from well calibrated correlations (Leps, 1973), from precedent on

similar dams or from large scale triaxial tests if required.

The dynamic resistance vector NW is represented by the force PQ in the force polygon which

corresponds to the minimum force required to mobilize the “reserve” friction angle (

sliding surface. The force, PQ, is the minimum force requir

direction that makes an angle of 90 degrees with respect to the direction of the resultant R. The

magnitude of the dynamic resistance NW, can be estimated from the triangle OPQ as:

Fig. 14 Force Polygon of

sin (

Thus, the yield acceleration Ng can be estimated as: Ng = sin (

3.3 Calculation of Seismic-Induced Displacements along Potential Failure Surface.

The method proposed by N. Newmark, 1965 to ap

seismic-induced displacements is based on the estimates of the relative motion of a rigid block with

respect to the ground triggered by acceleration pulses with a magnitude larger than the yield

acceleration.



induced acceleration, An, on any sliding mass on the slope can be estimated as proposed by

(Ambraseys and Sarma, 1967), and described above. If the maximum acceleration A

sliding mass is lower than Ng, then the slope is safe and no seismic-induced displacements will take

place. On the other hand, if the induced acceleration, An, on any of the sliding masses considered,


ude of the accumulated displacements that take place along the sliding surface during these

periods of time when the seismic-induced acceleration exceeds the “yield acceleration” are calculated

using the simplified method proposed by N. Newmark (1965), described below.



representative of the potential sliding mass in Fig. 13. In the force polygon of Fig. 14 the magnitude

and direction of the weight vector W are known. The mobilized friction angle,

maintain static equilibrium along the sliding surface is also known as and is equal to


known, and is determined by the friction angle, with respect to the normal, N, to the sliding plane.

, of well compacted rockfill materials, generally ranges between 45 to 50


similar dams or from large scale triaxial tests if required.

ic resistance vector NW is represented by the force PQ in the force polygon which

corresponds to the minimum force required to mobilize the “reserve” friction angle (

sliding surface. The force, PQ, is the minimum force required to close the force polygon and has a



Fig. 14 Force Polygon of Sliding Wedge

sin ( - α1) = �

= N

Thus, the yield acceleration Ng can be estimated as: Ng = sin ( - α1) g

Induced Displacements along Potential Failure Surface.

The method proposed by N. Newmark, 1965 to approximate the magnitude of the accumulated

induced displacements is based on the estimates of the relative motion of a rigid block with


8


, on any sliding mass on the slope can be estimated as proposed by

(Ambraseys and Sarma, 1967), and described above. If the maximum acceleration An on the potential

induced displacements will take

, on any of the sliding masses considered,


ude of the accumulated displacements that take place along the sliding surface during these

induced acceleration exceeds the “yield acceleration” are calculated



Fig. 13. In the force polygon of Fig. 14 the magnitude

and direction of the weight vector W are known. The mobilized friction angle, m, required to

maintain static equilibrium along the sliding surface is also known as and is equal to α1, the inclination


with respect to the normal, N, to the sliding plane.

, of well compacted rockfill materials, generally ranges between 45 to 50


ic resistance vector NW is represented by the force PQ in the force polygon which

corresponds to the minimum force required to mobilize the “reserve” friction angle ( – α1) along the

ed to close the force polygon and has a



Induced Displacements along Potential Failure Surface.

proximate the magnitude of the accumulated

induced displacements is based on the estimates of the relative motion of a rigid block with



9

The results of this method are summarized in Fig. 15, which provides a relationship between the ratio

N/A, the coefficients of the yield acceleration over the seismic-induced acceleration on the sliding

wedge, and the corresponding seismic-induced displacements of the wedge. A review of the results

shown in Fig. 15 indicates that the magnitude of the seismic-induced displacements of the sliding

wedge can be classified in three different groups depending on the magnitude of the N/A ratio. For

those cases where the N/A ratio is equal to or greater than 0.5, the seismic-induced displacements are

likely to be relatively low, less that 6 in., and would not have a significant effect on the embankment

behavior. For cases, where the N/A values range between 0.5 and 0.20 the seismic-induced

displacement can be significant, up to 3 ft (1m), but still acceptable if adequate free board has been

provided. The potential off set of the filter materials caused by these displacement levels can be

tolerated, and the strength of the embankment materials is not significantly reduced by the shear strain

imposed by these displacements. However, if the N/A ratio drops below 0.20, the resulting seismic-

induced displacements can be large, with a magnitude of several feet, and can trigger unacceptable

behavior, including loss of free board and overtopping, significant shear strength loss along the siding

surface, and/or unacceptable shear off sets within the filter layer. As a general design criteria, in

highly seismic zones, the N/A ratio should be maintained above 0.2.

Fig. 15 Standardized Displacement for Normalized Earthquakes, Unsymmetrical Resistance

(from Newmark, 1965)

The magnitude of the seismic-induced displacements for potential sliding wedges with N/A ratios

lower than 0.2 can be estimated as:

δ = 6 νmax 2/2 gN

where δ = seismic-induced displacements

νmax = peak particle velocity at the center of gravity of the sliding wedge, in

feet/sec

g = acceleration of gravity in feet/sec2; and

N = yield acceleration coefficient.


10

The magnitude of the peak particle velocity, νmax, can be estimated from the ratio of νmax/An = 150

cms/sec/g; where An is the seismic-induced sliding wedge acceleration. For example if the

acceleration, An, at the center of gravity of a shallow sliding wedge is equal to 0.7 g, the corresponding

peak particle velocity, νmax, would be equal to 3.88 ft/sec. If the sliding wedge has a yield acceleration

coefficient of 0.1 the seismic-induced displacements can be estimated as 11 ft.

4. DESIGN RECOMMENDATIONS

In highly seismic areas, the design criteria of large CFRD should include a minimum N/A ratio of

0.20. Because, the critical sliding wedges are located near the crest of the dam, special measurements

to achieve this criteria include:

a) Flatten the slopes in the upper sections of the dam.

The downstream slope near the crest should be flatter than the average downstream slope as shown in

Fig.16. The upstream slope at the upper sections of the dam should be flatter in case of reservoir

fluctuating with high frequency.

PARAPET

Zone 3BZone 3A

Zone 2B

Face Slab

Zone 1B

Zone 1A

Zone 2A

1.41.0

1.51.0

1.41.0

Flatter Slope

Depends of Dam height (25 - 30% H)

.

.

8+2% Hm

Zone 3C

Zone 4

DRAIN

1A COHESIONLESS SOIL - COMPACTED BY CONSTRUCTION EQUIPMENT

1B RANDOM - COMPACTED BY CONSTRUCTION EQUIPMENT

2A PROCESSED MATERIAL (Ø MAX. = 3

4 ") - MANUAL COMPACTION

2B PROCESSED MATERIAL (Ø MAX. = 3" - 4 ") 4 - 6 PASSES OF 12 Ton VIBRATORY ROLLER

3A SELECTED SMALL ROCK PLACED IN SAME LAYER THICKNESS AS ZONE 2

3B QUARRY RUN ROCKFILL, ABOUT 0,60m TO 0,80m LAYERS, 4 - 6 PASSES OF 12 Ton VIBRATORY ROLLER

3C QUARRY RUN ROCKFILL, ABOUT 0,80m TO 1,00m LAYERS, 4 - 6 PASSES OF 12 Ton VIBRATORY ROLLER

4 DOWNSTREAM ROCKFILL - PLACED ROCKFILL

Fig.16 Suggested seismic design for CFRDs to 300m high.

b) Widen the crest of the dam as the height increases.

Crest wide should be 8m minimum for dams up to 150m high. For dams higher than 150m to 300m

increase the crest width to : w = 8 + 0.02H where H: height of the dam in meters.

c) Provide a zoning of the dam as shown on Fig.16 The zone 3B of well graded rockfill materials should be extended at the upper portion of the dam (25 –

30%H) to the downstream slope. This criterion is being applied in some CFRDs already under

construction.


11

d) Provide good compaction to the rockfill

Specify compaction with vibratory rollers with a minimum of 12t over the cylinder and having a ratio

of static weight/cylinder width, equal or >5t/m. Restrict lift thicknesses to 0.8m and apply water with

monitors with a ratio of 200 l/m³as a minimum.

e) Provide a higher freeboard than conventional.

Besides the hydraulic requirements for the conventional calculation of freeboard an additional 0,3%H

should be added in areas of high seismicity.

f) Limit height of parapet walls

Height of parapet walls should be limited to 3 to 4 m high.

g) Provide external water stops with more capacity.

Besides the traditional copper water stop located internally at the perimetric joint, an external

corrugated or with additional capacity water stop should be designed

h) Reduce the slab width in zones of steep topography.

The traditional width of face slab lanes of 15m, should be split to 7,50m in steeper zones where

differential settlement can occur during heavy shaking.

i) Treat carefully construction joints between face slab stages. Experience has demonstrated that some gaps can be developed behind the struded curb under the

construction joint of the face slab stages. These gaps should be well grouted to avoid serious

displacements during heavy shaking.

j) Place compressible filler materials in the central longitudinal joints of the concrete

slab.

For narrow canyons where the A/H² < 4, experience has demonstrated that it is necessary to place

compressible fillers to mitigate compression stresses during the settlement caused by reservoir filling

and severe ground shaking. A=Concrete face area in m²; H=Height of the dam in meters.

REFERENCES

[1] N.M. Newmark, “Effects of Earthquakes on Dams of Embankments,” Geotechnique 15:140-

141, 156, 1965.

[2] N. N. Ambrasey and S. K. Sarma, “The Response of Earth Dams to Strong Earthquakes,”

Geotechnique 17:181-213, 1967.

[3] F.I. Makdisi and H.B. Seed, “Simplified Procedure for Estimating Dam and Embankment

Earthquake-Induced Deformations,” Journal of the Geotechnical Engineering Division, ASCE,

Vol. 104, No. GT7, July, 1978, pp. 849-867.

[4] J.B. Cooke and J.L. Sherard “CFRD Design, Construction and Performance” – ASCE

Symposium, Detroit, Michigan, 1985.

[5] L. Arrau, I. Ibarra, G. Noguera. “CFRD Design, Construction and Performance” CFRD book,

Performance of Cogoti Dam Under Seismic Loading, ASCE Symposium, Detroit, Michigan,

1985.

[6] Chen Shengshui et al, Analysis of Seismic Permanent Deformation of Rockfill Dam, Journal of

Nanjing Hydraulic Research Institute, 1990(3): 277


12

[7] Xu Zeping, Performance of the Zipingpu CFRD During the Wenchuan Earthquake, Hydropower

Dams, Issue three, 2009.

[8] M. Wieland, Chairman ICOLD Committee on Seismic Aspects of Dam Design – “CFRD’s in

Highly Seismic Regions” – Water Power and Dam Construction, April 2010.

[9] P.T. Cruz, B. Materón and M. Freitas, “Concrete Face Rockfill Dams” – Book, Released in the

Occasion of the 23rd

ICOLD Congress, Brasília, Brazil, May 2009.

[10] N. Matsumoto et als, Performance of the Ishibuchi CFRD During the Miyage Earthquake,

Hydropower & Dams, Issue One, 2011.

[11] Y.Yamaguchi et al, Preliminary Investigation of Dams Stricken by the Iwate-Miyagi Nairiku

Earthquake, 2008.Article from Google.

[12] K. Ishihara, Performance of Rockfill Dams During Recent Large Earthquakes, Fifth

International Conference on Recent Advances in Geotechnical Earthquake Engineering and Soil

Dynamics, San Diego, California, 2010.

[13] R.J. Marsal and L. Ramirez, “Performance of El Infiernillo Dam, 1963-1966,” Journal of the

Soil Mechanics and Foundations Division, ASCE, Vol. 93, No. SM4, July, 1967, pp. 265-298.

[14] D. Resendiz, M.P. Romo and E. Moreno, El Infiernillo and La Villita Dams: Seismic Behavior,

Journal of the Geotechnical Engineering Division, ASCE, Vol. 108, GT1, January 1982.

[15] T.M. Leps, “Flow Through Rockfill,” Embankment Dam Engineering, John Wiley and Sons,

1973.

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Documents

seismic design of cfrds

cfrd design

seismicinduced displacement

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