ONGOING SEISMIC SAFETY ASSESSMENT OF SWISS DAMS (*) R

ONGOING SEISMIC SAFETY ASSESSMENT OF SWISS DAMS (*)
R. PANDURI
Swiss Federal Office of Energy, Dam Safety Section, Switzerland
P. DROZ
Stucky Ltd., Renens, Switzerland
S. MALLA
Axpo AG, Baden, Switzerland
R. RADOGNA
Officine Idroelettriche della Maggia SA, Locarno, Switzerland
M. WIELAND
Pöyry Energy Ltd., Zurich, Switzerland
G. R. DARBRE
Swiss Federal Office of Energy, Commissioner for Dam Safety, Switzerland
1.
INTRODUCTION AND MOTIVATION
The largest Swiss dams were designed and built in the 50’s and 60’s.
These include the four highest Swiss dams of Grand Dixence (285 m), Mauvoisin
(250 m), Luzzone (225 m) and Contra (220 m). For all dams, the seismic safety
had been verified according to the state of knowledge at the time of design. Typical assumptions were e.g. a horizontal peak ground acceleration of 0.1 g independent of the site, the use of a pseudo-static approach, and the consideration of
the dynamically entrained water mass with the formula of Westergaard.
Since then, earthquake engineering has made important steps forward
concerning the state of knowledge. This comprises the earthquake hazard, the
analysis methodologies, and the underlying safety concept, to mention a few. So
there is the necessity to reassess the seismic safety of existing Swiss dams
based on the current state of knowledge and the current safety philosophy. The
Swiss guidelines prepared at the early 2000s define the technical basis for this
reassessment. They are largely a formalization of the engineering practice in
Switzerland.
The seismic safety assessments of Swiss dams according to the new
guidelines are being carried out since 2003. A seismic safety assessment for all
dams which fall under the validity of the ordinance should be available within 10
years of the publication of the guideline, i.e. till 2013. At present, about one third
of all seismic safety assessments have been delivered to the supervising authori(*)
Courantes vérifications aux séismes des barrages Suisses
ty. The next imminent step is thus to complete the ongoing assessment and validate the results. Where needed, safety measures for seismic retrofit will be mandated.
2.
2.1
TECHNICAL BASIS
SWISS GUIDELINES FOR THE EARTHQUAKE SAFETY ASSESSMENT OF DAMS
From 2000 to 2003, the supporting document for the earthquake safety assessment of Swiss dams has been prepared by a working group commissioned
by the dam supervising authority of the Swiss Confederation. This document is
still the valid basis for the verification of the seismic safety of dams [1].
The primary goal of the seismic safety assessment of dams is the protection of the population against loss of life and against property damage. Central to
the implementation is the underlying philosophy according to which the order of
magnitude of the risk shall be similar for all dams. This implies that the chosen
probability of occurrence of the safety evaluation earthquake (SEE) depends on
the consequences of a dam failure, which finally led to the specification of three
dam classes. In addition to the probability of occurrence, the allowed procedures
depend on the dam class too, permitting for smaller dams simplified methodologies in comparison to larger dams.
To fulfill the goal of the seismic safety assessment, the main (and sole) minimal requirement in the supporting document is that for the specified SSE, no
failure with uncontrolled release of reservoir water occurs. Consistent with this
goal, the authority requires a stringent safety assessment for the SEE only, and
not for a basis operation earthquake, the latter being left to the owner’s choice.
Based on the same philosophy, minor or even important damages to the dam or
to appurtenant structures which however do not jeopardize the integrity of the
dam or the operation of safety-relevant structures are explicitly allowed under the
SEE.
2.2
SEISMIC INPUT
Existing seismological groundwork at the time of the preparation of the
guidelines form the basis of the probability-based definition of the SEE, which is
given by a set of acceleration response spectra and effective peak accelerations
for various return periods. The normalized shapes of the response spectra correspond to those of Eurocode 8. They apply both for horizontal and vertical directions. The peak ground acceleration as scaling value is defined starting from
maps provided in the guidelines, depending on the dam class and thus on the
return period of the SEE. In effect, intensity maps are provided in the guidelines,
from which the horizontal peak ground acceleration is obtained by means of an
empirical intensity – acceleration relation. The vertical peak ground acceleration
is eventually taken as two thirds of the horizontal component.
Where time-history analyses are required, namely for the highest dam
class, at least three analysis runs with different sets of stochastic independent
time histories need to be done, each set consisting of two (2D) or three (3D)
components. For this purpose, sample normalized time histories which can be
used for the safety assessment have been provided by the safety authority.
2.3
MODELING REQUIREMENTS IN THE GUIDELINE
The supporting document for the seismic assessment of dams is composed
of different sections. After a general section and a section about the definition of
the seismic input at the site, the document defines the proposed methodology
depending on the dam types. Accordingly, the different methodology sections
deal with earth- and rockfill dams, with concrete and masonry dams as well as
with weirs.
The seismic safety assessment of earth- and rockfill dams is based on a
two-step analysis. At first, the stability of individual dam parts is systematically
evaluated. If the stability is not guaranteed and sliding occurs, a sliding analysis
is performed in a second step. It must be demonstrated that the overall dam stability is still guaranteed after the potential sliding. This implies in particular that a
sufficient freeboard remains and that the drainage and core layers continue to
fulfill their intended purpose. The modeling requirements depend on the dam
class, ranging from simplified stability analyses with an equivalent horizontal action to dynamic finite-element calculations.
In the verification of the seismic safety of concrete and masonry dams, both
a stress and a stability assessment need to be performed. In the stress analysis,
usually linear elastic, the stresses resulting from the combination of static and
dynamic loads are compared to the material strength. In case of overstressing, it
must be demonstrated that stress redistribution is possible and that no local instability endangering the structure will occur. In the stability analysis, overturning
and sliding of the whole structure are investigated. Here too, the modeling requirements, depending on the dam class, range from simple beam models with a
pseudo-static analysis to finite element models with a time history analysis.
Since weirs can encompass a large variety of structural systems and materials, it was not possible to prepare universally valid general recommendations as
for the other types of dams. This was however not deemed necessary, as a weir
usually consists of different types of structures for which general guidelines already exist in the other sections of the supporting document. In any case, the
fundamental philosophy of the guidelines must be considered.
3. EXAMPLES OF ONGOING ASSESSMENTS
3.1
EARTHQUAKE SAFETY ASSESSMENT OF MATTMARK ROCKFILL DAM
The 117 m high Mattmark dam constructed from 1960-1967 rests on a soil
layer with a maximum depth of 88 m (Fig. 1). It was the first large dam to be built
on top of a deep soil layer. Details about this unique embankment dam can be
found in Gilg ([2], [3]). The dam is situated in the highest seismic zone of Switzerland. In connection with a planned heightening of the reservoir level, a comprehensive seismic safety evaluation was carried out based on the draft of the new
Swiss earthquake guidelines for dams. The Mattmark study was the first practical
application of this document [4].
The dynamic analysis was performed using a two-dimensional finite element model of the dam-foundation system. The input motion consisted of three
different sets of spectrum-compatible artificial accelerograms with horizontal and
vertical peak ground accelerations of 0.42 g and 0.28 g, respectively, corresponding to the safety evaluation earthquake (SEE) with a return period of 10,000
years. Originally the dam was designed against earthquakes using a seismic
coefficient of 0.1, the critical load combination being rapid drawdown and earthquake (pseudo-static) acting in unfavorable direction.
The dynamic material properties were selected on the basis of the results
of static and dynamic laboratory tests and information available in the geotechnical literature [5]. The cyclic triaxial laboratory tests on two types of soils (cylinder with diameter of 30 cm and height of 60 cm) representing essentially the core
and filter material of Mattmark dam have demonstrated that the generation of
excess pore water pressures during cyclic loading after 50 cycles is not significant because the soils tested behaved dilatant. The residual deformations at the
end of cyclic loading can be substantial however, i.e. 5% or more.
The dynamic deformation analyses showed that shallow slides could move
2 m to 3 m during the SEE. For deeper slides, the maximum displacement would
be less than 80 cm. From the safety point of view, the deeper slides are more
critical, as they cut through the whole core, whereas the shallow slides do not.
The seismic settlements were estimated based on the reduction of the shear
stiffness of the various materials during the SEE.
For the seismic safety assessment of the dam the following performance
criteria were specified:
- The freeboard after the SEE shall be such that the water level in the reservoir
is below the top of the core. The freeboard remaining after the sliding displacements and the seismic settlements caused by the SEE is sufficient after
the earthquake.
- For the prevention of piping along slip surfaces at least 50% of the filter and
transition zones must still be functional after the earthquake. The core shall
exhibit ‘self-healing’ properties. The well-graded core material taken from a
moraine and the wide filter layers will prevent any piping effects even in the
case of sliding movements of 2 to 3 m.
In the calculation of the yield acceleration, the unfavorable effect of the vertical component was taken into account by applying a constant upward acceleration corresponding to the root-mean-square (r.m.s.) value of the average vertical
acceleration of the sliding mass. However, this leads to rather low yield accelerations and thus to an overestimate of the sliding movements of both the upstream
and downstream slopes and thus of the vertical deformations of the crest.
Fig. 1
View of Mattmark dam and cross-section with inclined core and grout curtain
Vue du barrage Mattmark et coupe avec noyau incliné et écran d'injection
3.2
EXAMPLES FOR GRAVITY DAMS
In the following we discuss three examples of increasing complexity of
seismic assessment of concrete gravity dams. The first example concerns the
Robiei gravity dam, built in the period 1964-67. The dam is divided into 21 blocks
of 17m width each. The main section has a height of 67.5m, while the base is
50m wide. Since the dam is a Class 1, there was performed a non-stationary 2D
linear elastic analysis. To this purpose there were generated 3 sets of 2 independent spectrum-compatible accelerograms each. The horizontal peak ground
acceleration is ahor,peak = 0.26g and the stationary duration of all time histories is at
least 18s. The fluid-structure interaction between the dam and reservoir is taken
into account by introducing additional masses in the structural system according
Westergaard. The area mostly stressed in traction of the dam in question is at the
toe of the upstream face, where the maximal tensile stresses (3.3Mpa) do find
place. However, the maximum tensile stress is always smaller than the dynamic
tensile strength of concrete (3.5MPa). On the other hand, the maximum compression stresses at the downstream face, at toe level, are far lower than the dynamic compressive strength of concrete (35.1MPa).
In addition to stress assessments, for gravity dams global stability of equilibrium must also be checked. The global stability was first evaluated by means of
a pseudo-static approach. Since the safety factor was less than 1 for both sliding
and overturning (mainly because of the severe hypothesis dictated by the guidelines of simultaneous acting of the horizontal and the vertical acceleration peaks)
at a later stage a non-stationary dynamic analysis (Newmark), considering the
sliding of the dam in the dynamic field, was applied. The same three pairs of accelerograms used in the previous stress-strain analysis were again applied. It is
to be observed that in the elastic analysis the contemporaneity of horizontal and
vertical acceleration peaks never happens because of the independency of artificially generated accelerograms. This situation penalizes seismic verification of
gravity dams compared to arch dams. Due to earthquake, the maximum displacement of Robiei dam would be slightly less than 10cm. Such a shift, not a
priori negligible, however, involves neither a danger to the structure itself, nor for
the security disposals such as bottom outlet and spillway, as they are external to
the dam body. The stability verification is successful even in presence of sliding
because an uncontrolled outflow of water is avoided.
Slightly more complex is the second example, concerning the dam Naret II,
with special regard to the assessment of strength of materials. While the compressive stresses are not critical, in two time-histories the admissible value of the
dynamic tensile strength of concrete was slightly exceeded (2.2>2.0Mpa). The
verification would not be satisfied in the first moment. However, analyzing the two
critical stress histories, it could be seen that the exceeding of the dynamic tensile
strength of concrete is modest and occurs only in two cycles. The seismic security is met according to the guidelines.
Finally we face the third example concerning the dam of Gries. This example is more complex because of the peculiarity of incorporating in its body hydro
mechanical equipment, namely the bottom outlet and the penstock of the hydroelectric power plant (Fig. 2).
Fig. 2
Gries dam – View and crown section
Barrage de Gries – Vue et coupe principale
As in the first example, all requirements of strength of material have been
met and are no further treated. The assessment of global stability, however, is of
interest. Also for this dam a non-stationary dynamic analysis, considering the
sliding of the dam, was necessary. Subsequent to the earthquake, the final displacement would be 1.3cm. Such a movement does not involve a danger for the
dam body. However, further investigations are necessary as to the penstock and
the bottom outlet.
For the penstock, a separated new FE model with beam elements was
built. The shortening of 1.3cm was imposed as upstream boundary condition.
The ideal stress according to von Mises is 162 MPa, admissible for the steel of
this pipe. The seismic safety is therefore satisfied.
The bottom outlet consists of a steel pipe DN1800mm drowned in concrete.
This pipe is tapered to a rectangular section 1000x1300mm where two rectangular gates (revision and exercise) reside. The guideline states that the security
organs must remain efficient after an earthquake, or that their function can be
restored quickly. By the FEM calculation model of the dam, along the gate guide,
there is a drift  = hor / hvert = 0.0001. The smallness of this relative motion ensures the permanence in the elastic range of the gate and of its guides. We must
also consider the possibility of manoeuvring the gates after the earthquake.
Thanks to the presence of a manual pump, allowing to drive even in absence of
electricity, and thanks to the possibility of a rapid assembly of flexible pipes
(which should be provided in all valve chambers), even in case of damage of the
steel hydraulic pipeline system, it is guaranteed to drive the gates rapidly. The
verification is therefore satisfied.
3.3
EXAMPLES FOR WEIRS
The classification of dams and the procedures which are to be used for
their seismic assessment applies also in the case of weirs. But in comparison to
large dams, their geometry may be much more complex and the hydromechanical structures play usually a major role regarding the integrity of the water retention system.
For the weirs of Class 1, a three-dimensional finite element model of the
dam, including appurtenant structures, gates and foundation is used. Calibration
and validation of the numerical model are carried out through comparisons between measured and calculated dam temperatures and displacements. Static
analyses of the dam are first conducted to evaluate stresses and displacements
under usual load cases, i.e. self weight, silt and sediments, hydrostatic pressure
and temperature gradient. Then, using site spectrum-compatible accelerograms,
dynamic analyses are run, in combination with initial static loads. Calculated
compressive and tensile stresses in the dam and appurtenant structures are
compared to dynamic concrete strength. The stability of the dam against sliding
and overturning is evaluated considering the maximal dynamic response. Finally,
seismic safety of hydromechanical structures and gates is evaluated considering
steel strength, jamming hazard and buckling of hydraulic jacks [6].
Verbois Dam, owned and operated by SIG, Geneva, is located on the
Rhone River downstream Geneva. The dam is 36 m high above foundation, for a
crest length of 427 m. The volume of the reservoir is 13.6 Mm3. The water retention structure consists of different parts, including a concrete gravity dam on the
left abutment with embankment shouldering, a gated weir part with 4 sluices 14
m wide each, a power plant part, an operation building part and a concrete gravity dam on the right abutment (Fig. 3).
Foundation is mostly sandstone. The maximal peak ground acceleration to
be considered for the seismic assessment is 0.27g. Three 3-D accelerograms
have been generated accordingly.
Fig. 3
Global model of Verbois Dam with crest and radial gates models
Modèle général du barrage de Verbois avec vannes clapet et segment
The various parts of the dam as well as the gates have been modeled independently. Their respective eigenfrequencies have been calculated for adjusting the damping parameters. The global model combines the various parts.
Joints between parts have been modeled considering linear elastic behavior with
high rigidity in the out-of-plane direction and frictional behavior in the plane. Tensions are analyzed a posteriori taking into account that the joints may open and
release the eventual tensile stresses [7]. The US/DS maximal displacements obtained are about 3 mm at the crest of the weir and HPP parts. They are of the
same order of magnitude transversally for the weir part and about 1.8 mm for the
HPP part. For the gravity parts, the US/DS maximal displacements are about 5
mm and 1.3 mm transversally. Crest gates present transverse maximal displacements of 2 mm but the US/DS and vertical displacements are ten times
larger. Radial gates show transverse maximal displacements of 1 mm, 4 mm
US/DS and 6 mm vertically.
The maximal tensile stresses are close to the dynamic tensile strength of
concrete (3.5 MPa); the compression stresses are far lower than the compressive
strength. The maximal tensile stress obtained in the steel structure of the crest
gates reaches 350 MPa (elastic threshold) in one of the girder and 250 MPa in
the radial gates. The safety with respect to buckling of the jacks is also confirmed. Comparing the maximal transverse displacements of the concrete structure of the weir with those of the gates and taking into account the initial spacing
indicates that the hazard of gate jamming can be discarded.
Finally, the global stability of the various parts is checked considering two
hypotheses (c' = 0 and ' =56° or c' = 500 kPa and ' =35°). A minimal factor of
safety of 1.07 is reached for the right bank concrete gravity dam considering the
first very conservative hypothesis, and 1.81 considering the second one. In conclusion, the safety of Verbois Dam regarding seismicity can be considered as
satisfied.
3.4
EXAMPLES FOR ARCH DAMS: ROGGIASCA
The 68 m high Roggiasca arch dam, which was completed and first impounded in 1965, is located in the canton of Grisons in south-east Switzerland.
With a crest thickness of 2.5 m and a maximum thickness at the base of 7.5 m
only (Fig. 4), the dam has a relatively high Lombardi slenderness coefficient of 21
[8]. The safety evaluation earthquake (SEE) with a return period of 10000 years
has a horizontal peak ground acceleration (PGA) of 0.17 g and a vertical PGA of
0.11 g. The three components of earthquake excitation were simulated by artificially-generated spectrum-compatible acceleration time histories.
Fig. 4
Central cross-section of Roggiasca arch dam and example time history of tensile
principal stress
Coupe central du barrage de Roggiasca et exemple de variation temporelle de la
contrainte de traction principale
The earthquake behavior of the dam was analyzed using a 3D finite element (FE) model, which comprised the dam, the foundation rock, the sediment
deposit on the upstream side and the soil filling on the downstream side. All the
dynamic calculations were performed using the FE Software ADINA [9].
The first three natural frequencies of the dam are 4.6, 5.4 and 7.6 Hz in the
empty reservoir condition and 2.9, 3.4 and 4.8 Hz in the full reservoir condition.
The seismic response is dominated by the third mode, which has the largest
modal mass participation factor in the along-stream direction. The largest accelerations and stresses due to the earthquake shaking occur at the center and the
two quarter points of the dam crest. The upstream sediment deposit and the
downstream soil filling were not found to play a significant role in the dynamic
behavior of the dam.
The largest compressive stress due to the combination of static and dynamic loads is about -10 MPa, which is not a problem for the dam concrete with a
static compressive strength of 46 MPa.
The linear-elastic analysis shows that the earthquake excitation produces
tensile principal stresses of up to about 6 MPa oriented in the arch direction in
the crest region (Fig. 4). Even when combined with the compressive stresses due
to the static loads (self-weight and hydrostatic pressure), the maximum tensile
stress is still about 4 MPa. If the thermal stresses in winter and effect of the ongoing chemical expansion of the dam concrete are also considered, the tensile
stresses would become even higher. The linear-elastically computed tensile
stresses for the earthquake load combination would exceed the dynamic tensile
strengths of the contraction and lift joints, which are substantially weaker in tension than the monolithic mass concrete. Thus, the vertical contraction joints
would open and cracks would develop at the horizontal lift joints, possibly leading
to formation of detached concrete blocks in the crest region. Such detached
blocks would be subjected to sliding and rocking motions during a strong earthquake.
The seismic stability of potential detached concrete blocks of various
heights in the crest region was analyzed using 2D models, one of which is illustrated in Fig. 5. In view of the geometric constraints in an arch dam, such a detached concrete block cannot move beyond the downstream face of the dam.
This was simulated in the numerical model by means of gap elements. The
cracked lift joint was represented in the model as contact surfaces. The base of
the detached block was subjected to the amplified acceleration at the crack level,
as obtained from the 3D linear-elastic analysis of the arch dam [10].
As the Roggiasca arch dam is rather thin, any detached block would be
quite slender. Such a detached block has a tendency to undergo only rocking
with virtually no sliding. However, even the rocking motion of the block is not very
large, causing dynamic crack openings of a few millimeters only at the base (Fig.
5). The detached blocks remain dynamically stable even when subjected to peak
horizontal accelerations about twice as large as the pseudo-static overturning
acceleration. This can be explained by the fact that an acceleration peak in an
excitation with a dominant frequency of about 5 to 8 Hz exceeds the pseudostatic overturning acceleration only for a very short time less than one-tenth of a
second, which is too short to result in any significant block rotation. A review of
literature on dynamic overturning of rigid blocks subjected to ground motion also
confirmed that acceleration peaks would have to be many times larger than the
pseudo-static overturning acceleration for a block with dimensions of the order of
a few meters to be toppled by a dynamic base excitation at such a frequency
([11], [12]).
Fig. 5
Model with gap elements and example time history of crack opening
Modèle avec des éléments à seule compression et exemple de variation temporelle pour l’ouverture d’une fissure
4. EXAMPLES OF UPGRADING PROJECTS FOR SEISMIC SAFETY
4.1
SEISMIC STRENGTHENING OF ILLSEE DAM
The 25 m high Illsee dam, located in the canton of Valais in south-west
Switzerland, is being currently rehabilitated and strengthened to meet the requirements of the Swiss guidelines for earthquake safety assessment of dams
and to mitigate the effects of the chemical expansion of the dam concrete. This
dam was originally built in 1926-27 and later heightened by 7 m in 1941-43. It
consists of a curved portion (arch-gravity dam) on the left side and a straight
gravity dam on the right side, as shown in Fig. 6. In the curved portion, a horizontal crack can be seen at elevation 2353 m a.s.l., which follows the joint between
the older concrete of 1926-27 and the newer concrete of 1941-43.
Fig. 6
Layout plan and developed elevation showing upstream face of Illsee dam
Plan et élévation développée du parement amont du barrage de Illsee
The seismic safety of the dam was checked for a safety evaluation earthquake (SEE) with a return period of 10000 years. The SEE excitation, which has
a horizontal PGA of 0.44 g and a vertical PGA of 0.29 g, was simulated by artificially-generated spectrum-compatible acceleration time histories. The stability of
the following two typical sections of the dam was investigated:
- In the curved portion, the highest cross-section representative of blocks 17
and 18 was analyzed. The stability checks were performed for the horizontal
crack at elevation 2353 m a.s.l. The arch action was not taken into account,
as the upper part of the arch-gravity dam does not have a proper rock abutment on the right side, where it is supported only by the adjoining block
(block 14) at the end of the straight portion. Moreover, the planned cutting of
slots in the dam to relieve stresses due to the chemical expansion of the
mass concrete will further weaken the arch action in the upper part of the
curved portion.
- In the straight portion of the dam, the stability of block 5 was analyzed at the
concrete-rock interface at the base. This is the most critical block in the
straight portion, as the dam base is nearly horizontal here. All the remaining
blocks have bases sloping significantly upwards in the downstream direction
and therefore possessing a greater resistance to sliding.
The uplift pressure was assumed to decrease linearly from 100% to 0%
over the dam thickness. The friction angle was taken as 45° and the cohesion
was ignored. Both the investigated sections have the minimum required factors of
safety of 1.50 against sliding and overturning for the normal static load combination. For the ice load combination, however, the factors of safety are less than the
required minimum value of 1.30. To meet the stability criteria for this load combination, blocks 5 (straight portion) and 17/18 (curved portion) would need anchor
forces of 410 and 490 kN/m length of dam, respectively.
The dynamic stability was first analyzed for forces obtained from the linearelastic earthquake analysis of 2D FE-models of the sections, whereby a damping
ratio of 10% was assumed and the hydrodynamic effect of the reservoir was simulated using added masses. In both the curved and straight portions, the factors of safety against sliding and overturning were found to be considerably
smaller than 1.0, which implied that the SEE shaking would cause sliding and
rocking motions.
The nonlinear seismic response of the dam was analyzed by modeling the
crack in each investigated section as contact surfaces (ADINA R & D, 2008). This
analysis indicated that the upper part of the curved portion of the dam above the
horizontal crack at elevation 2353 m a.s.l. would slide towards the downstream
side by 1 m to as much as 2 m and crack openings of up to about 300 mm would
occur due to the SEE ground motion. If the friction angle would be smaller than
45°, the sliding displacement would be even larger and the stability of the dam
could be endangered. Therefore, the curved portion of the dam will be strengthened by a combination of post-tensioned and passive anchors carrying a total
force of 1900 kN/m length of dam (Fig. 7), due to which the minimum factor of
safety increases to 1.10 for the SEE loading.
In the case of block 5 in the straight portion, the nonlinear analysis without
any anchors showed that the SEE shaking would cause sliding displacements in
the range of 22 to 60 cm and minor crack openings of 2 to 7 mm, which would
not endanger the stability of the dam. Therefore, an anchor force of 410 kN/m
length of dam will be provided in the straight portion to meet the stability requirement for the ice load combination only.
Fig. 7
Strengthening with anchors and time histories of factors of safety against sliding
and overturning of upper part of curved portion of dam above horizontal crack at
elevation 2353 m a.s.l. after installation of anchors
Renforcement avec des ancrages et variation temporelle des facteurs de sécurité
pour le glissement et basculement de la partie du barrage supérieure à 2353 m
In addition to the seismic strengthening with anchors, slots will be cut in the
dam to relieve the stresses due to the chemical expansion of the dam. Moreover,
new drainage boreholes will be drilled as shown in Fig. 7 to further enhance the
dam stability and the downstream face will be cleaned and resurfaced with shotcrete to repair the frost damage. Furthermore, the existing spillway will be replaced by a new spillway to increase the discharge capacity and satisfy the freeboard requirements.
4.2
SEISMIC STRENGTHENING OF MELCHSEE DAM
The following describes the upgrading of Melchsee dam, a loose material
dam built at the end of the 50s in the Swiss Alps in the canton Obwalden (operator is EWO, Kerns. The designer is IM Maggia Engineering Ltd, Locarno).
Melchsee dam is an earth dam 14 m high, whose hydraulic seal is guaranteed by
a central impervious core (Fig. 8).
Fig. 8
Melchsee dam – Typical cross section
Barrage de Melchsee – coupe typique
The dam has been designed in 1955. At that time merely an horizontal acceleration of 0.1g was considered acting on the dam, without any amplification.
According to current guidelines, however, the dam is subject not only to a peak
horizontal ground acceleration of 0.24g, which is amplified by the resonance of
the structure to 0.36g, but also to vertical acceleration of 0.16g.
The results of seismic assessment under the current provision of the existing dam has shown that there is no danger of liquefaction of dam body, no danger neither for the hydro mechanical equipment nor for the stability of the reservoir banks. Not even the upstream slope of the dam is at risk. However, the
downstream slope of the dam is at risk, when calculated according uncoupled
procedure of Makdisi-Seed with deformation of 43 cm. Therefore, in order to upgrade the existing structure to the new guidelines, there is a need of strengthening the downstream slope.
The mechanical characteristics of the reinforcement material must ensure
at least the following properties:  = 20kN/m3 and  = 30 °, usual values for loose
material. At the time of writing this article, we do not know the final mechanical
parameters of the reinforcement material nor the grading curve.
From the topography of the valley an amount of at least 5,500 m 3 of material is necessary for the upgrading works. A particular boundary condition very
favorable to the planned project of strengthening of the dam is the future construction along the side of the mountain of a new cable car for recreational purpose. As part of that project will be a considerable amount of excavated rock material, especially in the neighbourhood of future foundations of the pylons. The reuse of this material to reinforce the Melchsee dam is ideal for both projects.
Obviously, the existing dam monitoring instrumentation, consisting of two
piezometers and six geodesic measuring points, must be extended, but only to
the statically determinant part of the dam. Figure 9 shows the typical section of
the concession plan, where two new piezometers (one right in the impervious
core re-using a coring required for geotechnical characterization of the nucleus)
and new geodetic measurement points at the toe of the new embankment can be
seen. There must be also a number of inspection wells to control the drainage
system of water downstream, as the stagnation would decrease the static and
dynamic safety of the new dam.
Fig. 9
Melchsee dam – Strengthening and new monitoring system of the dam
Barrage de Melchsee – Renforcement et nouveau système d’auscultation
4.3
SEISMIC STRENGTHENING OF LES TOULES DAM
Les Toules double curvature arch dam is located in Switzerland close to the
southern border with Italy, in Canton of Valais. The owner of the dam is Forces
Motrices du Grand-St-Bernard (FGB). The dam is 86 m high and the crest across
the valley is 460 m long. The concrete volume is 235’000 m3. The ratio between
crest length and dam height is 5.35, which is far beyond common values for double curvature arch dams. The dam presents a particular design, with a slender
shape, a high vertical curvature toward downstream and no shear keys. An internal so-called Prepakt joint exists at the contact between the initial dam (1958)
and the heightened dam (1964). Les Toules dam was designed in the late 1950s
with the knowledge and practice of earthquake engineering of that time. Due to
limited knowledge on seismic hazard, most of the old dams in Switzerland have
been designed for a peak ground acceleration (PGA) of 0.10g.
Based on the dam classification, the PGA that should be considered is
0.33g which is 3.3 times higher than the original design earthquake of the dam.
To be more specific and relevant, a seismic hazard assessment was carried out
in order to determine a more precise PGA based on local faults and site geology
and select three proper recorded earthquakes corresponding to local conditions.
The seismic hazard study of the site was performed according to deterministic
and probabilistic approaches and resulted in a horizontal PGA of 0.28g and 0.19g
for the vertical component [13].
Static and dynamic analyses of the dam have been performed in order to
check its behaviour and assess its seismic resistance. The FE model and the
characteristics of the materials have been fixed after calibration of the displacements considering the past loading conditions (water level, temperature). The
dynamic calibration of the dam was carried out on the basis of natural frequencies and also mode shapes, both obtained by ambient and forced vibration tests
of the dam. The dynamic modulus of the dam concrete and the behavior of the
construction joints could be inferred from the approach and used for the timehistory dynamic analysis of the dam.
The static analysis showed relatively high vertical tensile stresses on the
upstream face of the dam for full reservoir load case, up to 5 MPa in winter. The
compressive stresses were also significant but acceptable with a maximum value
of 12 MPa for static load cases. The results of the dynamic analysis showed that
the dam would experience very high tensile stresses in case of a seismic event.
The maximum tensile stress obtained was about 12 MPa. Such high vertical tensions confirmed the necessity of the dam strengthening. In addition, high horizontal tensions were obtained in the central part of the upper arches, which could
trigger significant opening of the radial joints.
After having considered various options, it was found that reinforcing the
dam on its downstream face by addition of two lateral strengthening masses (abutment thickening) on each bank, together with shear columns in the central section of the arch was the optimal solution (Fig. 10), in particular since it does not
involve any work on the upstream face [13].
Fig. 10
Cross-sections of existing dam with strengthening solution - Left: Bloc 21 with
abutment thickening; Center: Shear keys at the crest; Right: Bloc 6 with abutment
thickening and a 10 m deepening into the rock foundation
Coupes type indiquant le renforcement – Gauche : épaississement de la culée;
Centre: clé de cisaillement; Droite: épaississement de la culée et approfondissement de la fondation de 10 m
In-depth static and dynamic analyses of the strengthened arch dam were
performed, showing that for the static cases, the high tensile stresses observed
in the upstream face of the existing dam are reduced by 45% for the critical load
case. The dynamic analysis confirmed that the increase of the dam rigidity by
thickening of the arches leads to increased values of natural frequencies. A comparison of the results obtained for the existing dam and for the strengthened dam
shows a reduction of the tensile stresses by 30% [14].
The construction started in 2008 was completed in time in 2011. Usual
techniques related to dam technology have been implemented, but needed to be
adapted: strengthening work type, hydropower always in operation during construction, harsh weather conditions.
5. CONCLUSIONS AND NEXT STEPS
Initially, owners and engineers viewed the seismic safety guidelines with
skepticism, as they feared that only few dams could fulfill the requirements set
forward in the guidelines. After several assessments performed for dams belonging to all classes, it becomes clear that the safety requirements will be met without the need for remedial measures in all but a few cases, some of which are
presented in this paper.
These exceptions, where the seismic safety requirements defined in the
guidelines cannot be met, are mostly cases for which the state and behavior had
already been questioned under normal operational conditions or that have a particular structural configuration. In other words, dams that satisfy current construction and safety standards for normal operating conditions were generally also
found to satisfy the requirements related to seismic actions.
This should actually not come as a surprise, the primary structural purpose
of a dam being to transfer the large horizontal component of hydrostatic forces
into the abutments and foundation. The lessons learned from large earthquake
events worldwide also confirm the excellent behavior of dams under earthquake
loading.
The Swiss guidelines proved to be a good basis for the ongoing seismic
safety assessments. For a future revision of the guidelines, one major concern
will be the consideration of new findings for the earthquake hazard in Switzerland.
ACKNOWLEDGEMENTS
The authors wish to acknowledge the Owners of the various dams mentioned in this paper, in particular Services Industriels de Genève (Verbois) and
Forces Motrices du Grand-St Bernard (Les Toules). Special acknowledgements
are also due to A. Wohnlich and O. Müller for their commitment in Les Toules
Dam rehabilitation project and A. Mellal and A. Tzenkov detailed studies of the
Verbois Dam seismic analysis.
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Supporting Document for the Earthquake Safety Evaluation of Swiss
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American Society of Civil Engineers, pp.215-248, 1974.
GILG B. Mattmark, Swiss Dams, Monitoring and Maintenance, Edition for
15th International Congress on Large Dams, Swiss National Committee on
Large Dams, Lausanne, pp. 89-95, 1985.
All Swiss Guidelines for the Safety of Dams together with their base documents can be
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tests for seismic safety assessment of an earth dam, Proc. 16th Int. Conf.
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arch dam subjected to strong ground shaking. Proceedings of the 1st European Conference on Earthquake Engineering and Seismology
(ECEES), Geneva, Switzerland, Sept. 3-8, 2006.
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WOHNLICH A., MÜLLER O.; Strengthening of Les Toules arch dam,
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SUMMARY
Since the design of the largest Swiss dams in the 50’s and 60’s, earthquake engineering has made important steps forward concerning the state of
knowledge. So there is the necessity to reassess the seismic safety of existing
Swiss dams.
The seismic safety assessments of Swiss dams according to the new
guidelines prepared at the early 2000s are being carried out since 2003. Initially,
owners and engineers viewed the seismic safety guidelines with skepticism, as
they feared that only few dams could fulfill the requirements set forward in the
guidelines. After several assessments performed, it becomes clear that the safety
requirements will be met without the need for remedial measures in all but a few
cases. This should actually not come as a surprise, the primary structural purpose of a dam being to transfer the large horizontal component of hydrostatic
forces into the abutments and foundation. The lessons learned from large earthquake events worldwide confirm the excellent behavior of dams under earthquake loading.
The Swiss guidelines proved to be a good basis for the ongoing seismic
safety assessments. For a future revision of the guidelines, one major concern
will be the consideration of new findings for the earthquake hazard in Switzerland.
Les plus grands barrages Suisses ont été construits dans les années 50 et
60. Dès lors, les connaissances dans le domaine de l’ingénierie sismique ont
évoluées. Ainsi, les vérifications des barrages Suisses aux séismes doivent être
réévaluées.
Les vérifications selon les nouvelles directives préparées au début des années 2000 sont en cours dès 2003. Au début, les propriétaires et ingénieurs affrontaient les nouvelles directives avec scepticisme, comme ils craignaient que
seulement une petite partie des barrages puissent satisfaire les critères des directives. Après des divers vérifications, il se trouve que de mesures de confortement ne sont nécessaires que dans une minorité des cases. En effet cela n’est
pas une grosse surprise comme le but principal des barrages est de transférer
les grands components horizontaux dus aux forces hydrostatiques dans les appuis et les fondations. Les expériences faites au niveau mondial à l’issue de
grands événements sismiques confirment le comportement excellent des barrages sous des charges sismiques.
Les directives Suisses ont confirmé être une bonne base pour les vérifications courantes. Pour une future révision des directives, un des points principaux
sera la prise en considération de nouveaux connaissances quant à l’aléa sismique en Suisse.