This is not the document you are looking for? Use the search form below to find more!

Report home > Manual & Guide

Ductility Design of Foundations of Highway Bridge Abutments

0.00 (0 votes)
Document Description
civil engineering reference manual to design Foundations of Highway bridge abutments that has introduced the seismic design in its calculation
File Details
  • Added: May, 10th 2010
  • Reads: 1102
  • Downloads: 116
  • File size: 769.78kb
  • Pages: 15
  • Tags: ductility design, earthquake, highway analysis, highway bridge, reference, seismic design
  • content preview
Submitter
Embed Code:

Add New Comment




Related Documents

Seismic design of earth-retaining structures and foundations

by: mersada, 9 pages

Earthquake-resistant design of earth retaining structures like retaining walls, earth dams and foundations are very important problems to minimize the devastating effect of earthquake hazards. In ...

DOE - Design Of Experiments For Dummies – A Beginner’s Guide

by: heidi, 10 pages

Design Of Experiments For Dummies – A Beginner’s Guide By Willy Vandenbrande | Published: 21 Feb 07 Despite all the efforts by specialists in quality and statistics, Design ...

DOE - Design Of Experiments QC & Taguchi Methods

by: karin, 2 pages

DESIGN OF EXPERIMENTS, QC AND TAGUCHI METHODS Now, more than ever before, processing costs and problems of repeatability can stall new composite programs right at the ...

Primer for Design of Commercial Buildings to Mitigate Terrorist ...

by: radenka, 6 pages

It is impractical to design a civilian structure to remain undamaged from a large explosion. The protective objectives are therefore related to the type of building and its function. For an office, ...

The Evolving Design of Protest: Interview series

by: ugen, 19 pages

An interview in the series titled 'The Evolving Design of Protest' with Derek Rasmussen, an aboriginal rights activist. This is part of a series of interviews conducted by Urgyen Badheytsang for his ...

The Evolving Design of Protest: Interview series

by: ugen, 15 pages

An interview in the series titled 'The Evolving Design of Protest' with John, a former Greenpeace activist and current teacher at OCAD. This is part of a series of interviews conducted by Urgyen ...

Ui Design Of Maemo 5 Apps Digia

by: colin, 19 pages

Ui Design Of Maemo 5 Apps Digia

DESIGN OF EXPERIMENTS BASED FORCE MODELING OF THE FACE GRINDING PROCESS

by: shinta, 8 pages

A grinding force model is developed to predict the forces during the face grinding of cast iron and aluminum alloy 319. Design of experiments methods are used to create a response ...

Design of CMU Common Lisp

by: manualzon, 122 pages

ebook about Artificial Intelligence on LISP: Design of CMU Common Lisp.pdf, useful ebook for newbie or intermediate LISP programming

Computer Analysis & Reinforced Concrete Design of Beams

by: manualzon, 129 pages

ebook design computer analysis and reinforced design of beams, for structure engineering and civil engineering project with computer aided design

Content Preview
DUCTILITY DESIGN OF FOUNDATIONS OF HIGHWAY BRIDGE
ABUTMENTS
Masahiro SHIRATO∗, Jiro FUKUI†, and Junichi KOSEKI‡
Abstract
Ductility design of foundations of abutments against severe earthquakes has been
newly introduced in the Specifications for Highway Bridges in Japan. This paper describes
the summary of the ductility design method as well as its background. One of the features
of this method is the adoption of a new procedure to evaluate seismic active earth pressure
that is applicable up to high seismic loads based on the modified Mononobe-Okabe method.
We also conduct back-analyses of the case histories of abutments including foundations
damaged in the past earthquakes following the ductility design method proposed here. The
results from the back-analyses confirm the applicability of the ductility design method.
Introduction
Right after the 1995 Hyogo-ken Nanbu earthquake (Kobe earthquake) in Japan, the
Specifications for Highway Bridges and commentaries were revised in 1996[1, 2]. The
specification has introduced ductility design of foundations of piers as standard seismic
design against severe and rare scale earthquakes. It should be noted that the aseismic design
of piers based on the ductility design itself was not new, but had been already introduced
before the 1995 Hyogo-ken Nanbu earthquake against large earthquakes which correspond
to type I motions of level 2 earthquake motions defined in detail later.
Nevertheless, abutments and their foundations have been exception in applying the
ductility design even in the 1996 version of the specification, and the specification still
required to check that they behave elastically against level 1 earthquake motions which are
small to middle (or frequent scale) earthquakes as it had demanded in the past. The reasons
for this judgment are the following. First, because of the lack of knowledge on a proper
evaluation method of their performances, including the seismic active earth pressure at high
seismic loads, it is difficult to establish any verification methods of seismic performance of
abutments and their foundations against severe earthquakes. Second, there has been no case
history in Japan where the damage to abutments or their foundations causes unseating of
girders. Third, a bridge abutment is a structure resisting against the earth pressure exerted
from the backfill at any moment, and thus it would be just pushed forward by the seismic
∗Research Engineer, Foundation Engineering Research Team, Public Works Research Institute
†Team Leader, ditto.
‡Associate Professor, Institute of Industrial Science, the University of Tokyo

earth pressure even if it suffers residual displacements during earthquakes, which would
not directly result in the unseating of supporting girders.
However it is needed to develop an aseismic design method of abutments and their
foundations against large earthquakes, which controls their damage within an acceptable
extent to recover the service of road network as early as possible after the earthquakes.
After publishing our 1996 version of the specifications, we have constructed a database of
past earthquake damage to abutments or abutment foundations in Japan, and have studied
on the assessment methods which can distinguish the different damage extents for the case
histories in our database. Separately, Koseki et al.[4] have recently proposed an evalua-
tion method of active earth pressure at high seismic loads, and it enabled us to establish
a computational method to assess seismic performance of abutment foundations. Conse-
quently, we could newly introduce the ductility design of abutment foundations into the
latest version of the Specifications for Highway Bridges revised in March 2002. The de-
sign method is based on the preceding frame work of the ductility design method of bridge
pier foundations, while some modifications are implemented to it in order to reflect the
above-mentioned peculiar mechanical characteristics of foundations of abutments. This
paper describes the verification procedure and the process of the development of the new
ductility design method of foundations of abutments.
Fundamentals of seismic design of abutments and their foundations
A two stage aseismic design procedure is applied on the basis of the combinations
of the level 1 and level 2 earthquake motions and corresponding requirements of struc-
tural performance. The level 1 earthquake motions are likely to strike a structure once or
twice during the expected service period of the structure. Their peak amplitudes are small
to medium, and are around 0.2 G at the ground surface. On the other hand, the level 2
earthquake motions are extremely strong, but very unlikely to strike a structure during its
service period. The level 2 earthquake motions include two types of motions. One is called
‘type I’ motions which are generated at plate boundaries in the ocean. Their peak ampli-
tudes at the ground level are smaller than the other one although the type I motions have
longer durations. Their peak amplitudes are about 0.3 G to 0.4 G at the ground surface.
The other motions are inland strike type motions called ‘type II’ which are produced by a
fault located near the site; for example the 1995 Hyogo-ken Nanbu earthquake is catego-
rized in this type. The type II motions have high intensities but short duration. Their peak
amplitudes are set at 0.6 G to 0.8 G at the ground surface, based on the acceleration records
on the ground surface observed in the 1995 Hyogo-ken Nanbu earthquake.
In the new aseismic design procedure, all abutments and their foundations are veri-
fied against the level 1 earthquake motions so that they would cause no structural damage.
The abutment and its foundation shall remain in the elastic response region.

Next their seismic performance against the level 2 earthquake motions is checked.
However this process is not necessarily applied to all highway abutments and their founda-
tions. For only the foundations to be constructed on/in liquefiable ground, the verification
against the level 2 earthquake motions shall be conducted. In this case, the foundation
shall have the necessary strength and ductility to fulfill the structural requirements as fol-
lows: the damage to the foundation shall be limited within a level where it can be repaired
with reasonable cost and shall not cause a state where rescue operation of the bridge is not
available. These requirements refer to those for pier foundations which undergo effects of
liquefaction. Consequently, the ductility design against the level 2 earthquake motions is
applied only to deep foundations in liquefiable ground.
For abutment bodies above their foundations, verification may be omitted against
the level 2 earthquake motions even in the new specification as well as in the former speci-
fications. This is because an abutment which satisfies the performance requirement for the
level 1 earthquake motions is considered to posses enough strength and ductility against
the level 2 ones. This can be also supposed by the past damage case histories. It should be
noted that some structural details need to be specified in order to supplement skipping the
assessment of the seismic performance against the level 2 earthquake motions. The new
specification requires to arrange lateral confining reinforcements in the abutment body so
as to improve its ductility. It also demands to include the same amount of longitudinal
reinforcements on the front side of the cross section of the abutment body as those on its
back side in order to prevent excessive bending failure and diagonal tensile failure caused
by contacting of the abutment with the deck.
Survey of case histories of damage to abutments
We made a list of damaged abutments including foundations from the past damage
reports of recent large earthquakes in Japan. We selected typical cases from the list as is
summarized in Table 1. Table 1 exhibits contents of damage, height of abutment from the
bottom of pile cap, bearing condition, and equivalent thickness of liquefiable soil strata,
which is explained below, for each case. Note that we here picked up only case histo-
ries damaged by the type II earthquake motions to simplify the comparison among these
cases, because the behavior of abutments, foundations, and sub-soil layers would highly
depend on characteristics of seismic motions, e.g. the number of cycles, duration time, and
the intensity of earthquakes. All abutments that were finally picked up are supported by
group-pile foundation. Note also that pile foundations are employed for about a half of
the highway bridge foundations in Japan, in particular those are constructed in deep soft or
deep loose soil layers.
We judged the rank of damage based on the contents of damage for each case. In

Table 1 List of past damage case histories of abutments and computed nonlinear response of their
pile foundations
Case
Earth-
Rank
Damage
Height
bearing
Equivalent
Response
quake*
of
of
condition
thickness of
ductility
damage**
abutment
liquefiable
factor
(m)
soil strata
HE (m)
A
1
4
- spalling of concrete around anchor bolts of
9.8
move
6.8
3.2
bearings
- movement of abutment (10 cm)
B
1
4
- tilting of abutment, shear collapse of wing-
5
move
19.7
4.2
wall
- slumping of backfill (1.0 m)
C
2
3
- tilting of abutment
3.3
move
18
not
- cracks in parapet wall
yield
- outward movement of foundation (11 cm at
the top of the foundation)
D
1
3
- outward movement of abutment
5.5
fix
13.4
1.8
- slumping of backfill
E
1
2
- slumping of backfill (20 cm)
8.5
move
3
not
- collapse of bearings
yield
F
1
2
- damage of abutment
8.8
move
3.2
1.4
- cracks in the surface of approach road
G
1
2
- cracks in parapet wall
10.3
move
6.9
1.1
- bump at the connection part between abut-
ment and backfill
H
1
2
- cracks of abutment
7
move
8
2
- collapse of side-blocks of bearings
- exchange of bearings in repair works
- slight slumping of backfill
I
1
2
- slumping of backfill
6.5
fix
1
1.2
J
1
2
- shrinkage of spacing of expansion joint
12.3
fix
0
2.9
- excessive movement of bearing
- bump in backfill
K
1
1
- a little spalling of concrete of abutment
6.6
fix
11.9
1.9
L
1
1
- shrinkage of spacing of expansion joint
7.5
move
4.7
1
- cracks in parapet wall
- bump in backfill
M
1
1
- shrinkage of spacing of expansion joint
6
fix
0.5
1.3
* ‘1’: the 1995 Hyogo-ken Nanbu earthquake, ‘2’: the 2000 Western Tottori earthquake.
** Refer to Table 2
doing so, we set a four-tier criterion for the damage rank emphasizing whether the service
of the bridge was available or not right after the earthquake. The larger the number of the
rank, the more disastrous the damage. The relation among the ranks, serviceability and
repairability, and details of damage is displayed in Table 2. The serviceability indicates
the effect of the damage to the abutment or its foundation on the service or function of
the bridge right after the earthquake, and the repairability represents whether the complete
recover of the abutment and its foundation was possible by repair work with reasonable
cost or not.
Equivalent thickness of liquefiable soil strata HE is defined by
HE = H∗ + H
1
2

Table 2 Categorization of the degree of damage
Rank
of
1
2
3
4
damage
Degree of
slight
medium to large
severe
damage
Service-
Fully operational
Operational
with
No operation tem-
Impossible
ability
some
restrictions
porarily while doing
w.r.t
weight
of
emergency counter-
vehicles and speed
measure works**
limit
Repair-
Easy*
Possible with minor
Possible with major
Impossible (Recon-
ability
repair works
repair works
struction)
Typical
- shrinkage of spac-
- slumping of back-
- horizontal move-
- excessive horizon-
damage
ing
of
expansion
fill
ment or rotation of
tal movement or ex-
contents
joint
abutment
cessive rotation of
abutment
- cracks of parapet
- cracks of structural
- excessive slumping
- collapse of struc-
wall
members
of backfill
tural members
- collapse of parapet
wall
* e.g., with fixing slight cracks
** e.g., operational with some restrictions after constructing temporary bents
H∗ = 1.5H
1
FL1 + 1.0HFL2 + 0.5HFL3
(0 m ≤ z ≤ 10 m),
H∗ = 1.0H
2
FL1 + 0.5HFL2
(10 m ≤ z ≤ 20 m),
where z is the depth of soil layer from the ground level, H∗ is the thickness of liquefiable
1
soil strata estimated for strata in between z = 0 m and z = 10 m, and H∗ is the one estimated
2
for strata in between z = 10 m and z = 20 m, where the origin of the ground level GL is set
at the bottom of pile cap. HFL1 is the sum of the thicknesses of the soil strata for which the
values of factor of safety against liquefaction FL are assessed to be less than or equal to 0.6,
HFL2 is the sum of those with 0.6 < FL ≤ 0.8, and HFL3 is the sum of those with FL > 0.8.
This index has been proposed as one of indices in a past guideline for the assessment of
earthquake resistance of existing bridges. The equivalent thickness involves the effects that
the shallower the location of the liquefiable soil stratum is, the larger the impact of the
liquefaction of that stratum is on the horizontal bearing capacity of the foundation against
loads acting on its top, and that the smaller the value of F L is, the larger the loss of soil
stiffness is. In this study, we estimated the values of F L by following the procedures for the
type II seismic motions described in the currently effective specification[5].
An example of the cases with the damage rank 4 is shown below. The case ‘B’ in
Table 1 is the abutments of Higashi Uozaki bridge which collapsed by the 1995 Hyogo-ken

constraint by contact with deck
tensile
earth
force
pressure
residual horizontal
displacement
loss of
subgrade reaction
liquefied soil strata
Figure 1 Schematic mechanism of damage to
Photo 1 Failure of abutments of Higashi
abutments of Higashi Uozaki bridge
Uozaki bridge[3]
Nanbu earthquake[3]. This bridge passed over a canal, and had a total length of 64.8 m with
three spans. The abutments were cantilever-type walls with pile foundations consisting of
H-shaped steel piles. Photo 1 shows the failure of the A1 abutment. The A1 abutment
rotated backward and large cracks widespread on the front face of the abutment. Besides,
in the A2 abutment wall, shear collapse with widely opening cracks on the front face was
also observed. Based on the observation that the backfill of the abutments slumped, we
estimated that the abutments suffered residual horizontal displacement. The subsoil layers
consist of loose or soft soils, namely slime, sand, silty sand, and clay with SPT-N blow
counts of zero to six till 14.7 m deep from the river bed. Sand boils caused by liquefaction
were observed, and the revetment suffered outward lateral movement in the direction to the
canal. As is illustrated in Figure 1, the backward rotation of the abutment accompanied
with a lot of cracks on its front face may have occurred as follows:
1) the liquefaction of the subsoil layers resulted in the shortage of horizontal subgrade
reaction to support the foundation;
2) the foundation was pushed outward by the earth pressure acting on the back face of
the abutment and the inertial force of the abutment itself;
3) the decks constrained the above lateral movement at the top of the abutment;
4) the foundation and the lower part of the abutment continued to move laterally with
the above constraint, and a large tensile force by bending was mobilized at the front
face of the abutment.
The cause of damage to ‘A’ abutment in Table 1 which is categorized into the rank 4 is

also the lateral movement of foundation. It was possibly triggered by loss of horizontal
subgrade reaction induced by liquefaction of subsoil layers.
Figure 2 shows the relation between the damage rank and the equivalent thickness
of liquefiable soil strata for the abutments listed in Table 1. Although no unique relation
except for some positive correlation can be found, this figure implies that there is a pos-
sibility of severe damage (the rank 4) if the equivalent thickness of liquefiable soil strata
exceeds five meters. These case histories suggest that liquefaction of subsoil layers is the
primary cause of damage to abutments and their foundations.
Based on the damage mechanism estimated above, it is expected that we can pre-
vent abutments from severe damage as long as the possible displacement of the foundation
of an abutment during an earthquake is well-controlled even though liquefaction of the
subsoil layers may occur. Therefore the ductility design of abutments is applied to only
their foundations, and the specification requires us to carry out it when the subsoil layers
are considered to be liquefiable.
Seismic loads employed in the ductility design method of foundations of abutments
In the ductility design method, effects of earthquakes are modeled with pseudo-
static loads, namely seismic active earth pressure acting on the back face of the backfill
located on footing, inertial force acting on the backfill, abutment itself (wall), and footing,
and horizontal reaction from deck as shown in Figure 3. It would be reasonable to evaluate
the intensities of these pseudo-static loads based on the characteristics of their dynamic
responses considering the effects of liquefaction. However the quantitative evaluation of
these effects is not well established. Thereby, we set practically the intensities of these
seismically induced loads as follows. Note that we do not take the vertical component of
seismic motions into account in accordance with the customary practice.
The design horizontal seismic coefficient for evaluating seismic active earth pres-
sure and the inertial force of backfill located on footing is obtained as cAkhg, where khg is
the design horizontal seismic coefficient at ground surface, and cA is the modification factor
for the amplification of acceleration in the backfill as well as in the adjacent embankment.
For simplicity, we assume that cA = 1 here. In Figure 3, khA denotes the design horizontal
seismic coefficient to determine the inertial forces of abutment and footing. We also as-
sume that khA = cAkhg here, because the behavior of abutments during earthquakes would
be dominated by that of the backfill and embankment. The coefficient η to be multiplied
by horizontal reaction force from deck in Figure 3 represents the effect of phase differ-
ence between the horizontal response of superstructure and that of embankment. However,
we adopt that η = 1 here for simplicity, and assume the direction of the horizontal reac-
tion force from deck to be forward to result in more severe load conditions for abutment

η× reaction
4
force from
deck
cAkhg
HE = 5 m
e
c
failure
Akhg
k
3
hA
plane
damag
of
2
Rank
khA
earth pressure
1
0
10
20
Equivalent thickness of liquefiable
soil strata HE (m)
Figure 2 Comparison between equivalent thick-
ness of liquefiable soil strata and the
Figure 3 Schematic figure of numerical model
rank of damage
foundations in practical design.
Seismic active earth pressure
The Mononobe-Okabe method (denoted herein as “M-O method”) is well-known
and widely accepted in a lot of design specifications or manuals as an evaluation method
of seismic active earth pressure. Examples of seismic active earth pressure coefficients ob-
tained by the M-O method are shown in Figure 4a with dotted lines, where kh is horizontal
seismic coefficient; KEA is seismic active earth pressure coefficient; φ is the internal friction
angle of embankment soil; δ is the friction angle at the interface between backfill located
on footing and embankment; and we assume the surface of embankment to be horizontal
and the back face of backfill on footing to be vertical. Designers usually employ φ = 30
to 40 degrees as the internal friction angle of embankment soil in the evaluation of earth
pressure. These values of φ are equivalent to residual strength of dense or relatively hard
soils in general. As their representative value, φ = 35 degrees was employed in Figure 4.
The levels of khg, the design horizontal seismic coefficient, at the ground surface are 0.6
to 0.8 in the ductility design, while it can be seen from Figure 4a that the values of KEA
with φ = 35 degrees cannot be obtained at such levels. Moreover, we examine the ratio
of failure zone length L as defined in Figure 4b to height H of abutment from the bottom
of footing. From the viewpoint of design practice, the failure zone length L indicates the
zone where settlement of road connecting to bridge may occur. The relations between L/H
and kh obtained by the M-O method are plotted with dotted lines in Figure 4b. When the
value of kh reaches 0.6 to 0.8, the values of L/H predicted by the M-O method with φ = 35

KEA (a)
L/H
Embankment
(b)
L
Interface
modified
Backfill
4
2
M-O
M-O method
φ
H
peak = 50◦
δ
φ = 50◦
failure
φres = 35◦
δ = 35◦/2
plane
failure plane
resultant force
of earth pressure
c
2
1
c
M-O method
M-O method
φ = 35◦
φ = 35◦
δ = 35◦/2
b
δ =
present
35◦/2
M-O method
modified M-O method
a
b
φ = 50◦
φpeak = 50◦, φres = 35◦
a
δ = 35◦/2
0
0
0
0.5
1
0
0.5
1
kh
kh
Figure 4 Seismic active earth pressure and failure zone of embankment (a, b, c: formation of first,
second, and third failure planes, respectively, in modified M-O method)
degrees become unrealistically large.
Now we recall the behavior of dense sand obtained by an element test (plane strain
compression test). Firstly the soil element is subjected to the process of strain hardening;
and then the element comes to mobilize the peak strength, and a shear band (or failure
plane) starts to emerge progressively; after that, strain softening process evolves, accom-
panying with the development of the shear band; finally the residual strength is mobilized.
Since backfill on footing of abutment and embankment soils are well-compacted, it would
be reasonable to take this behavior into the estimation of seismic active earth pressure. Ac-
cording to the above behavior, one can consider the following mechanism on the formation
of failure planes in the embankment soil and the mobilization of active earth pressure:
• the angle of failure plane is associated with the peak internal friction angle of em-
bankment soil, because the peak strength is mobilized in the soil elements when the
shear banding starts, and
• the value of active earth pressure is associated with the residual internal friction an-
gle, since the postpeak reduction of internal friction angle evolves after the formation
of shear band (or along the failure plane that has been already formed).
Based on the above hypotheses, the modified Mononobe-Okabe method has been proposed
by Koseki et al.[4]
Figure 4a also displays an example of the values of KEA calculated by the modified
M-O method with dash lines, where we assume the peak internal friction angle φpeak to be
50 degrees and the residual one φres to be 35 degrees; we also assume that the first failure
plain is formed at kh = 0; and we set the values of δ to be φpeak/2 until the formation of the

first failure plane (at kh = 0) and to be φres/2 after that, as the value of δ at the interface
between soils has been empirically set to be φ/2 in applying the original M-O method. In
the modified M-O method, after the formation of the first failure plane, the same failure
plane continues to be activated unlike the result from the original M-O method, because the
peak strength will be mobilized along any other failure plane in the embankment excluding
the first failure plane along which the residual strength is mobilized. However, when the
seismic coefficient kh is increased to a certain level, another failure plane will be activated,
because it gives the extremum of seismic active earth pressure, and at this level the second
failure plane is formed. The third, fourth, and other consecutive failure planes will be
formed in a similar manner.
The modified M-O method has a couple of advantages in comparison with the orig-
inal M-O method as can be seen in Figure 4. One is that the modified M-O method can
provide active earth pressure coefficient at large earthquake loads, even though the values
of kh get to 0.6 to 0.8 which are considered in the seismic design against the level 2 earth-
quakes. Another advantage is that the modified M-O method provides considerably smaller
and much realistic size of active failure zone in embankment than the conventional method
yields. In not only model test results on retaining structures but also case histories of actual
railway retaining walls damaged in the 1995 Hyogo-ken Nanbu earthquake, the observed
angle of failure planes formed in backfill cannot be explained without introducing the peak
strength of soils, as has been reported by Koseki et al.[7, 8] Consequently, the modified
M-O method has been introduced into the new specification.
We, however, have given some engineering judgments to the original paper by
Koseki et al.[4] in order to make the modified M-O method more suitable for practical
use. In the modified M-O method, we have to estimate the level of kh at the formation of
the first failure plane. Since, at this moment, no reasonable method to evaluate it has been
established, we give zero value to the level of kh at which the first failure plane emerges.
This is because a failure plane can be formed due to a slight displacement of foundation
under the working load conditions. Next, we assume that no more failure planes appear
after the second failure plane is activated. This is because when the third failure plane
appears, the embankment should have already deformed largely, and it is considered that
the application of limit equilibrium approaches would be difficult at such a level. Thirdly,
although we can derive both the first and second failure planes, we take only the second
failure plane into account in estimating the seismic active earth pressure coefficient KEA
values for all the values of kh, because the second failure plane would be usually active
at the ground acceleration level of 0.6 to 0.8 G corresponding to the level 2 earthquake
motions. From the viewpoint of practical design, this assumption is reasonable, since this
gives us the values of KEA close to those which have been used in the past design practice
at the ground acceleration levels of around 0.2 G corresponding to the level 1 earthquake
motions. Hence the dimensions of abutments and their foundations can become similar to

Download
Ductility Design of Foundations of Highway Bridge Abutments

 

 

Your download will begin in a moment.
If it doesn't, click here to try again.

Share Ductility Design of Foundations of Highway Bridge Abutments to:

Insert your wordpress URL:

example:

http://myblog.wordpress.com/
or
http://myblog.com/

Share Ductility Design of Foundations of Highway Bridge Abutments as:

From:

To:

Share Ductility Design of Foundations of Highway Bridge Abutments.

Enter two words as shown below. If you cannot read the words, click the refresh icon.

loading

Share Ductility Design of Foundations of Highway Bridge Abutments as:

Copy html code above and paste to your web page.

loading