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Figure 1. A typical cross-section of the GRS RW having a
FHR facing for Yamanote Line over Chuo Line in Tokyo,
constructed during 1995–2000.
1 INTRODUCTION
1.1 GRS retaining walls
For last about 25 years, several innovative rein-
forced soil structures were developed and construct-
ed as retaining walls (RWs), bridge abutments etc.
for transportations (i.e., railways, including high-
speed train lines, and roads) and others. In the late
1990’s, geosynthetic-reinforced soil (GRS) retaining
wall (RW) with staged-constructed full-height rigid
(FHR) facing was developed (Tatsuoka et al., 1997,
2007). Fig. 1 shows a typical case. This GRS RW is
characterized by staged-construction (Fig. 2a): i.e.,
firstly a vertical GRS RW is constructed with a help
of gravel-filled bags placed at the shoulder of each
soil layer. After major deformation of the supporting
ground and backfill has taken place, a full-height
rigid (FHR) facing (i.e., a lightly steel-reinforced
concrete layer) is constructed by casting-in-place
concrete on the wall face wrapped-around with
geosynthetic reinforcement (Fig. 2b). Therefore, suf-
ficiently high connection strength between the fac-
ing and the reinforcement layers can develop, which
results in high earth pressure on the back of the fac-
ing, therefore, a high stability of the backfill.
This new type GRS RW has been constructed at
more than 850 sites (Fig. 3a) and the total wall
length is now about 130 km (Fig. 3b) mainly for
GRS structures recently developed and constructed for railways and
roads in Japan
F. Tatsuoka Tokyo University of Science, Chiba, Japan
M. Tateyama Railway Technical Research Institute, Tokyo, Japan
J. Koseki
Institute of Industrial Science, University of Tokyo, Tokyo, Japan
Page limit 20
ABSTRACT: The construction of permanent geosynthetic-reinforced soil (GRS) retaining walls (RWs) with a
staged-constructed full-height rigid facing for railways, including high-speed train lines, roads and others
started about twenty five years ago in Japan. The total wall length is now more than 130 km, replacing con-
ventional gravity-type or cantilever steel-reinforced concrete (RC) RWs and steel-reinforced soil RWs. Many
were also constructed to reconstruct conventional type RWs and embankments that collapsed during recent
earthquakes, heavy rains, floods and storms. It is proposed to construct GRS coastal dykes with lightly steel-
reinforced concrete facings connected to geosynthetic reinforcement layers as tsunami barriers. By taking ad-
vantage of this GRS RW technology, a number of bridge abutments comprising geosynthetic-reinforced back-
fill were developed. The latest version for new construction, the GRS integral bridge, comprises a continuous
girder integrated to a pair of RC facing, not using bearings, while the backfill is reinforced with geosynthetic
reinforcement layers connected to the facings. Its high cost-effectiveness with a very high seismic stability is
described. The first prototype was constructed 2011 for a new high-speed train line. It is proposed to construct
GRS integral bridges and GR approach fills to restore conventional type bridges that collapsed by tsunami of
the 2011 Great East Japan Earthquake Disaster and also to newly construct many for new high-speed train
lines. Based on the GRS integral bridge technology, a new method to reinforce existing old conventional type
bridges was developed; the backfill is stabilized by large-diameter nails connected to the abutments and then
the girder is integrated to the abutments: i.e., nail-reinforced soil (NRS) integrated bridges.
railways and also for roads and other facilities. Any
problematic case, including poor performance dur-
ing severe rains and earthquakes, has not been re-
ported. This type of GRS RW has totally replaced
conventional RW types (i.e., gravity type and canti-
lever RC RWs) and steel strip-reinforced soil RWs
with discrete panel facing for new construction of
RWs for railways.
a)
b)
Figure 2. Staged construction of GRS RW: a) construction
steps; and b) details of connection between the facing and the
reinforced backfill (Tatsuoka et al., 1997, 2007).
The full acceptance of this GRS RW technology
by railway engineers is due to a very high cost-
effectiveness with a high-seismic stability as validat-
ed by high performance during the 1995 Kobe
Earthquake (Tatsuoka et al., 1998). This feature was
re-confirmed by high performance of a number of
GRS RWs of this type that had been constructed for
railway including a high-speed train line (Tohoku
Shinkan-sen) during the 2011 Great East Japan
Earthquake Disaster. On the other hand, a great
number of conventional type RWs and embankments
on level ground and slopes and in water-collecting
places were fully collapsed during these and other
earthquakes and heavy rains. Many of them were re-
constructed to GRS RWs of this type. The first
coastal GRS RW, which is about 10 m high facing
Pacific Ocean, was constructed for a length of about
1.0 km in 2010, replacing a gravity type RW that ful-
ly collapsed by scouring in the supporting ground
during a storm in the summer of 2007 (Tatsuoka &
Tateyama, 2012). Based on this technology, coastal
dykes and railway and road embankments having
nearly vertical wall faces or sloped slopes consisting
of geosynthetic-reinforced backfill with continuous
facing connected to the reinforcement layers as tsu-
nami barriers have been proposed.
In this paper, the characteristic features of this
GRS RW technology, typical case histories and les-
sons obtained from their behavior, in particular those
during earthquakes and heavy rains, are described.
a)
b)
Figure 3. a) Locations; and b) annual and cumulative lengths of
GRS RWs with FHR facing (as of March 2011).
1.2 GRS integral and NRS integrated bridges
The second category of innovative GRS structure is
a series of new type bridge systems developed based
on the technology of GRS-RW with FHR facing, de-veloped to alleviate several serious problems with
conventional type bridges (explained later). During
the first half of 1990’s, a number of GRS bridge
abutments were constructed with the girder placed
on a pair of sill beams, via a pair of fixed and mova-
ble bearings, placed immediately behind the facing
on the crest of the reinforced backfill. To alleviate
the problems of long-term settlement and unstable
behavior during earthquakes of the sill beams, this
abutment type was modified by placing the girder on
the crest of facings via a pair of bearings (Tatsuoka
et al., 2005). Finally, a more innovative bridge sys-
tem was developed, called the GRS integral bridge,
which comprises an integral bridge (with the girder
integrated to a pair of RC facings without using
bearings) and the backfill reinforced with
geosynthetic reinforcement layers connected to the
5)5) Completion of
wrapped-around wall
4)4) Second layer3)3) Backfilling & compaction
2) Placing geosynthetic &
gravel gabions
Gravel gabionGeosynthetic
1) Leveling pad for facing
Drain hole
6)6) Casting-in-place
RC facing
Lift = 30 cm
5)5) Completion of
wrapped-around wall
4)4) Second layer3)3) Backfilling & compaction
2) Placing geosynthetic &
gravel gabions
Gravel gabionGeosynthetic
1) Leveling pad for facing
Drain hole
6)6) Casting-in-place
RC facing
5)5) Completion of
wrapped-around wall
5)5) Completion of
wrapped-around wall
4)4) Second layer4)4) Second layer3)3) Backfilling & compaction3)3) Backfilling & compaction
2) Placing geosynthetic &
gravel gabions
Gravel gabionGeosynthetic
2) Placing geosynthetic &
gravel gabions
Gravel gabionGeosynthetic
1) Leveling pad for facing
Drain hole
1) Leveling pad for facing
Drain hole
6)6) Casting-in-place
RC facing
6)6) Casting-in-place
RC facing
Lift = 30 cm
30
Annual w
all
length
(km
)
10
0
20
Tota
l (
km
)
1995 20001990 2005
1995 Kobe EarthquakeTotal
Annual
Restart of construction of
new bullet train lines
2004 Niigata-ken-Chuetsu EQ
2010
140
40
120
100
80
60
130 km
2011 Great
East Japan
EQ
facings. The GRS integral bridge is much more sta-
ble against seismic loads and cyclic lateral dis-
placements at the top of the facing caused by sea-
sonal thermal deformation of the girder. From these
features and because of no use of bearings, this
bridge is much more cost-effective than the other
bridge types described above.
Referring to high cost-effectiveness with high per-
formance of GRS integral bridge, a new method to
reinforce existing old conventional type bridges was
developed: the backfill is stabilized by large-
diameter nails connected to the abutments, then by
integrating the girder to the abutments.
2 GRS RETAINING WALLS
2.1 Characteristic features
The three major characteristic features of the GRS
RW system (Figs. 1 & 2) are summarised below:
(1) Staged construction procedure (Fig. 2a): The first
advantage of the staged construction is that the con-
nection between the reinforcement and the facing is
not damaged by differential settlement between the
facing and the backfill during wall construction.
Then, construction on relatively compressible sub-
soil without using heavy piles for the facing and/or
using relatively deformable backfill (e.g., on-site soil
with a high fines content) becomes possible. Sec-
ondly, good compaction immediately at the back of
the wall face becomes possible, allowing sufficient
outward movements at the wall face mobilizing suf-
ficient tensile forces in the reinforcement during
wall construction.
A firm connection between the RC facing and the
geosynthetic-reinforced backfill can be ensured as
follows (Fig. 2b). Firstly, fresh concrete can easily
enter the inside of the gravel-filled soil bags
wrapped-around with geogrid reinforcement through
the aperture of the geogrid. Secondly, extra water
from fresh concrete is absorbed by gravel filling the
soil bags, which reduces negative effects of the
bleeding phenomenon of concrete. The soil bags
function as a temporary but stable facing during
construction, resisting against earth pressure gener-
ated by compaction works and further backfilling at
higher levels. Then, backfill-compaction becomes
more effective. The soil bags also function as a drain
after construction and as a buffer protecting the con-
nection between the FHR facing and the reinforce-
ment against relative displacement if it takes place
after construction. Moreover, to construct a conven-
tional cantilever RC RW, concrete forms on both
sides of the facing and its propping are necessary
and they become more costly at an increasing rate
with an increase in the wall height. With this type of
GRS RW (Fig. 2a), on the other hand, only an exter-
nal concrete form anchored in the backfill is used
without using an external propping is necessary,
while not using an internal concrete form (Fig. 2b).
Figure 4. Effects of firm connection between the reinforcement
and the facing (Tatsuoka, 1992).
(2) Use of a FHR facing constructed by If the wall face is loosely wrapped-around with geosynthetic re-inforcement without using soil bags, or their equiva-lent, at the shoulder of respective soil layers, or if the reinforcement layers are not connected to a rigid facing, no or only very low tensile forces are acti-vated at the connection between the facing and the reinforcement (Fig. 4a). Then, no or only very small earth pressure is mobilized at the wall face, which results in no significant lateral confining pressure in the active zone. This results in low stiffness and low strength of the active zone, which may lead to intol-erably large deformation of the active zone. On the other hand, with this new GRS RW system, the soil bags function as a temporary facing structure and high earth pressure can be activated at the wall face already before placing a FHR facing (Fig. 4b). As the wrapping-around geosynthetic reinforcement at the wall face is buried in the fresh concrete layer (i.e., the facing), eventually the reinforcement layers are firmly connected to the FHR facing. Then, the earth pressure that has been activated on the tempo-rary facing structure comprising soil bags is eventu-ally transferred to the FHR facing. That is, relatively large earth pressure, similar to the active earth pres-sure that develops in the unreinforced backfill, is ac-tivated to the FHR facing, which results in high con-fining pressure in the active zone, thus high stiffness and strength of the active zone. Then, high perfor-mance of the wall can be ensured.
A conventional type RW is a cantilever structure that resists against the active earth pressure from the unreinforced backfill (Fig. 5a). Therefore, large in-ternal moment and shear force is mobilized inside the facing while large overturning moment and lat-eral thrust force develops at the base of the facing. Thus, a pile foundation is usually used. These disad-vantages become more serious at an increasing rate with an increase in the wall height. Relatively large earth pressure, similar to the one activated on the conventional type RW, may be activated on the back
Unstable active zone Very stable active zone
Reinforcement
Connected
No connection strength High connection strength
High tensile forceLow tensile force
High confining
pressureLow confining
pressure
Unstable active zone Very stable active zone
Reinforcement
Connected
No connection strength High connection strength
High tensile forceLow tensile force
High confining
pressureLow confining
pressure
of the FHR facing of the new type GRS RW. De-spite the above, as the FHR facing behaves as a con-tinuous beam supported by reinforcement layers at many elevations with a small span, typically 30 cm, small forces are mobilised inside the facing structure (Fig. 5b). Hence, the facing structure becomes much simpler and lighter than conventional cantilever RWs. Besides, the overturning moment and lateral thrust force activated at the facing base becomes small, which makes unnecessary the use of a pile foundation in usual cases. Consequently, the GRS RW having a stage-constructed FHR facing becomes much more cost-effective (i.e., much lower construc-tion cost, much speedy construction using much lighter construction machines) than the conventional type cantilever RC RW.
a) b)
Figure 5. a) conventional type RW; and b) GRS RW with
FHR facing (Tatsuoka, 1992, Tatsuoka et a., 1997).
(2) The use of relevant reinforcement type: A poly-
mer geogrid is used for cohesionless soil to ensure
good interlocking and a composite of non-woven
and woven geotextiles for high-water content cohe-
sive soils to facilitate both drainage and tensile rein-
forcement. Then, low-quality on-site soil can be
used as the backfill if necessary.
(3) The use of relatively short reinforcement: The
length of geosynthetic reinforcement necessary for a
sufficient stability of GRS RW having staged con-
structed FHR facing is relatively short when com-
pared to metal strip reinforcement used for walls
having discrete panel facing. This is because: 1) the
anchorage length of planar geosynthetic reinforce-
ment to resist against the tensile load similar to the
tensile rupture strength of reinforcement is much
shorter; 2) a FHR facing prevents the occurrence of
local failure in the reinforced backfill zone by not al-
lowing the development of failure planes passing
through the wall face at an intermediate height; and
3) local failure in the facing, which may take place
with discrete panel or block facing and may result in
the failure of the whole wall, is difficult to take place
with FHR facing. Factors 2) and 3) become more
important when concentrated load is applied on the
top of, or immediately behind, the facing.
Figure 6. A GRS RW having FHR facing at Tanata, Kobe city;
a) immediately after construction; and b) one week after the
1995 Kobe Earthquake (Tatsuoka et al., 1997, 1998).
Figure 7. Railway embankment that collapsed during the
2004 Niigata-ken Chuetsu Earthquake and its reconstruction to
a GRW RW having FHR facing (Morishima et al., 2005).
2.2 Collapse of RWs and embankments by earthquakes and their reconstruction
Numerous embankments and conventional type RWs collapsed by earthquakes in the past. On the other hand, high performance of a GRS RW having stage-constructed FHR facing during the 1995 Kobe Earthquake (Fig. 6) validated its high-seismic stabil-ity. Many gently sloped embankments and conven-tional type RWs that collapsed by that and subse-quent earthquakes were reconstructed to GRS RWs of this type (Tatsuoka et al., 1977, 1998; 2007; Koseki et al., 2006, 2008; Koseki, 2012). Typically, three railway embankments supported by gravity type RWs on slope collapsed during the 2004 Nii-gata-ken Chuetsu Earthquake and they were recon-structed to GRS RWs of this type (Fig. 7). The adop-tion of this technology was due to not only much lower construction cost and much higher stability (in
Earth
pressure
Stress concentration
Earth
pressure
Reinforcement
8 July 1992a)
24 Jan. 1995b) 24 Jan. 1995b)
Silt rock
Sand rock
1:2.
0
13
.18
m
1:4 (V:H)
After remedy
work
Railway
(Jo-etsu line)
Shinano riiver
Gravel-filled steel wire
mesh basket
Rock bolt
After remedy work:
GRS-RW with a FHR facing; slope: 1:0.3 (V:H);
height= 13.2 m, vertical spacing of geogrid= 30 cm
Before failure: sand backfill including
round-shaped gravel on sedimentary soft
rock (weathered, more at shallow places)
After
failureFailed
gravity RW
Silt rock
Sand rock
1:2.
0
13
.18
m
1:4 (V:H)
After remedy
work
Railway
(Jo-etsu line)
Shinano riiver
Gravel-filled steel wire
mesh basket
Rock bolt
After remedy work:
GRS-RW with a FHR facing; slope: 1:0.3 (V:H);
height= 13.2 m, vertical spacing of geogrid= 30 cm
Before failure: sand backfill including
round-shaped gravel on sedimentary soft
rock (weathered, more at shallow places)
After
failureFailed
gravity RW
particular for these soil structures on steep slopes) but also much faster construction resulting from a significant reduction of earthwork when compared to the original gently sloped embankment with a gravity type RW.
Based on vast and very serious damage to railway structures during the 1995 Kobe Earthquake, the seismic design codes for railway structures were substantially revised (Tatsuoka et al., 2010a). The revised codes for soil structures have following sev-eral new concepts and procedures: 1) Introduction of very high design seismic loads
(i.e., level 2) and three ranks of required seismic performance.
2) Recommendation for the use of geosynthetic-reinforced soil structures.
3) Evaluation of seismic performance based on re-sidual displacement.
4) Use of peak and residual shear strengths with well compacted backfill, while ignoring apparent cohesion due to suction assuming that it fully disappears during heavy rains.
5) Design based on the limit equilibrium stability analysis.
6) An emphasis of good backfill compaction and good drainage.
7) No creep reduction factor applied to obtain the rupture strength of geosynthetic reinforcement in seismic design.
Figure 8. Locations and numbers of GRS RWs having FHR
facing for railways constructed before the 2011 Great East Ja-
pan Earthquake Disaster (Tateyama, 2012).
The 2011 Great East Japan Earthquake Disaster,
which took place 11th
March, is the most disastrous earthquake in Japan after the World War II. The damage was a combination of those from earthquake motions and the accompanying great tsunami. A great number of old embankments and RWs that
were not designed and constructed following the current seismic design standard collapsed. In com-parison, a number of GRS RWs of this type that had been designed based on the revised seismic design code described above and constructed in the affected areas of this earthquake performed very well (Fig. 8),
a)
b)
c)
Figure 9. a) Collapse of a wing RW (with a masonry facing
of a bridge abutment at site No. 11 in Fig. 8 (Nagamachi, Sen-
dai for Tohoku Freight line); and b) & c) its reconstruction to a
GRS RW (by the courtesy of the East Japan Railway Co.).
Several conventional type RWs and embankments that collapsed were reconstructed to GRS RWs of this type (Fig. 9). A very fast construction was one of the important advantages of this technology also in this case. In particular, the railway was re-opened at a restricted speed before constructing a FHR fac-
Morioka
Hananomaki
Ichino-
seki
Sendai
Epicenter
Fukushima
Tohoku Shinkansen
(high speed train)
Tsukuba Express (rapid
commuters’ train)
Tohoku Shinkansen
Tohoku Line
Tohoku Line
Tazawako Line
Senseki Line
Yard for Tsukuba Express
Number of GRS walls
Aomori: 58
Iwate: 23
Miyagi: 10
Akita: 1
Yamagata: 2
Fukushima: 1
Ibaraki 1
(total) (96)
Locations of GRS RWs
investigated in details
Tokyo
←Nagamatchi Higashi Sendai→
Irrigation channel
ing (Fig. 9b). Fig. 10 shows one of the three em-bankments that collapsed during an earthquake in-duced one day after the Great East Japan Earthquake Disaster and reconstructed to GRS RWs of this type.
Figure 10. One of the three embankments between Yokokura
and Morinomiya stations, Iiyama Line, that collapsed during
the Nagano-Niigata Border Earthquake and reconstructed to
GRS RWs (by the courtesy of the East Japan Railway Co.)
2.3 Collapse of RWs and embankments by heavy rains, floods and storms and reconstruction
Numerous embankments and conventional type RWs collapsed also by heavy rains, floods and storms in the past. GRS RWs (Fig. 2) were also con-structed to replace a number of embankments that were eroded and washed away by over-flowing flood water (Tatsuoka et al., 1997, 2007). Numerous embankments for roads and railways retained by gravity-type RWs along rivers and seashores col-lapsed by floods and storms, usually triggered by over-turning failure of the RWs caused by scouring in the supporting ground (Fig. 11a). The backfill is largely eroded resulting in the stop of the function of a railway or a road. This type of collapse is easy to take place, because the conventional type RW is a cantilever structure of which the stability is fully controlled by the bearing capacity at the bottom of the RW.
On the other hand, GRS-RWs with a FHR facing is not such a cantilever structure as above, therefore, much more stable against scouring in the supporting ground (Fig. 11b). It is particularly important that the major part of the backfill can survive even if the supporting ground is scoured. A high embankment retained by a masonry gravity-type RW at the lower part on the left bank of Agano river, Niigata Prefec-ture, for West Ban-Etsu line (a railway of East Japan Railway) was collapsed by a flood 30th July 2011 by the mechanism illustrated in Fig. 11a. It was recon-structed to an about 9.4 m-high and 50 m-long GRS RW with a FHR facing (Fig. 12).
a)
b)
Figure 11. a) Collapse of conventional type RW by scouring in
the subsoil (the numbers show the sequence of events); and b)
improved performance of GRS-RW with FHR facing.
Figure 12. Reconstruction of RW on the left bank of Agano
river, Niigata Prefecture, for West Ban-Etsu line to a GRS-RW
(by the courtesy of the East Japan Railway Co.).
Fig. 13 shows another case of this type of failure of a gravity-type RW for a length of about 1.5 km along seashore facing the Pacific Ocean, triggered by scouring in the supporting ground by strong ocean waves during a typhoon No. 9, 29
th Aug.
2007. The wall was reconstructed to a GRS RW with FHR facing (Figs. 13b & c). The FHR facing has a strong resistance against sea wave actions dur-ing storms, while the wall could be stable even if the supporting ground is scoured to some extent. On the other hand, discrete panel facing is not relevant in such a case as this, because the loss of the stability of a single or several panel(s) by scouring in the subsoil and/or erosion of the backfill from joints be-tween adjacent panels may easily lead to the failure of the whole wall.
≒6.0
5,0
2.0
1.0
After failure (before
reconstruction)
Mattress basket
Gravel
Stabilized soil
Welded
steel fabric
Slope before failure
Railway track
(Iiyama line) [All unit in m]
≒14.0
1. Scouring
3. Collapse of
embankment
2. Over-turning of RW
River bed
Flood
Flood
Scouring
GRS-RWs with a FHR facing has a
high resistance against scouring
(All units in m)
8.5722.808
0.3
Crusher run
(C-40)
Basal geogrid
(Td= 60 kN/m)
Short geogrid
(Td= 30 kN/m)
3.49
Ground improvement by
cement-mixing 5.0
3.0
2.0
Level pad
(concrete in
steel-mesh
cage
9.3
6
Design high water level
(38.14 m)
0.7
2.0
Ordinary low
water level
(26.08)
Long geogrid
(Td= 30 kN/m)
Height from the
base (m) =
13.1
12.9
11.5
8.9
4.51.0
a)
b)
c)
Figure 13. Seawall for Seisho by-pass of National Road No.
1 in Kanagawa Prefecture, southwest of Tokyo: a) collapse for
a length of about 1.5 km by Typhoon No. 9, 29th Aug. 2007; b)
a typical cross-section of GRS RW; and c) GRS RW under
construction (a & b: by the courtesy of the Ministry of Land,
Infrastructure, Transport and Tourism).
3 NEW BRIDGE TYPES
3.1 Problems with conventional type bridges
A conventional type bridge usually comprises a single simple-supported girder supported by a pair of abutments via fixed (or hinged) and moveable bearings, or multiple simple-supported girders supported by a pair of abutments and a single or multiple pier(s) via multiple sets of bearings. The approach fill is usually unrein-forced backfill. The abutment may be a gravity structure (unreinforced concrete or masonry) or a RC structure. The conventional type bridge has a number of drawbacks as described below (Fig. 14a).
a)
b) Figure 14. a) Several major problems with conventional type
bridges; and b) development of new bridge types for new con-
struction and reinforcement (Tatsuoka et al., 2008a, b; 2009).
Firstly, as the abutment is a cantilever structure
retaining unreinforced backfill, the earth pressure
induces large internal forces as well as large
thrust forces and overturning moment at its bot-
tom. This problem becomes more serious when
designed against severe earthquake loads. There-
fore, usually the abutment becomes massive and a
pile foundation becomes necessary more at an in-
creasing rate with an increase in the abutment
height. Secondly, although only very small
movement is allowed with the abutments, the
backfill are constructed after the abutments have
been completed. Hence, when constructed on
thick soft ground, many long piles may become
necessary to prevent any small movement of the
abutments caused by the earth pressure as well as
settlement and lateral flow in the subsoil. Large
negative friction may develop along the piles.
Thirdly, the construction and long-term mainte-
nance of the bearings and the connections be-
tween simple-supported girders are generally
costly. Fourthly, the bearings and the backfill are
the weakest elements against seismic loads. The
girder may dislodge at a moveable bearing by se-
vere seismic loads. A significant bump may be
formed behind the abutment by long-term settle-
Soil bags
Full-height
rigid facing
Geogrid Concrete cover
Scour protection
Selected
backfill (RC40)
Crushed gravel
(RC40)
Tie rod
Head
Steel
sheet pile
10 March 2010
Pacific
Ocean
Casting-in-place of
concrete for FHR
facing
(7) Long-term service: 7a. settlement by self-weight & traffic load; &
7b. large deformation by seismic load
(1) Piles
Supporting
ground
(3) Backfill
(4a) Earth pressure
(static & dynamic)
(4a) Displacement due to
earth pressure
(4b) Ground settlement and
lateral flow due to the weight
of backfill, and associated
negative effects on the piles
(i.e., negative friction &
bending).
(2)
abutment
(6) Girder (deck)
(5) Shoes (or bearing)
(fixed or movable);
5a. Long-term maintenance
5b. Low seismic stability
Numbers (1) – (7):
event sequence.
(7) Long-term service: 7a. settlement by self-weight & traffic load; &
7b. large deformation by seismic load
(1) Piles
Supporting
ground
(3) Backfill
(4a) Earth pressure
(static & dynamic)
(4a) Displacement due to
earth pressure
(4b) Ground settlement and
lateral flow due to the weight
of backfill, and associated
negative effects on the piles
(i.e., negative friction &
bending).
(2)
abutment
(6) Girder (deck)
(5) Shoes (or bearing)
(fixed or movable);
5a. Long-term maintenance
5b. Low seismic stability
Numbers (1) – (7):
event sequence.
Conventional
GRS RW Integral
GRS integral NRS integrated
Solving problems with backfill
Solving problems with girder and facing
Rehabilitation of existing conventional type bridges
Structural integration
Nailing
Combined
ment of the backfill due to its self weight, traffic
loads and seismic loads.
3.2 Integral bridge and GRS RW bridge
To alleviate the problems with the conventional type bridge (Fig. 14a), several new bridge types have been proposed (Fig. 14b), which are either for new construction or for reinforcement of existing old bridges. For new construction, the integral bridge comprises the girder structurally integrated to the facings without using bearings. This bridge type was developed to alleviate problems with the structural part (i.e., the girder and facings) of the conventional type bridge. This bridge type is now widely used for roads in the UK, the USA and Canada due mainly to a lower cost for construction and maintenance result-ing from no use of bearings and the use of a contin-uous girder.
However, as the backfill is not reinforced, thus not integrated to the abutments (or facings), the backfill and the structural part do not help each oth-er. Hence, this bridge type cannot alleviate some old problems with unreinforced backfill (described above). Moreover, as the girder is integrated to the abutments, seasonal thermal expansion and contrac-tion of the girder results in cyclic lateral displace-ments at the top of the abutments, which may result in: i) development of high passive earth pressure on the back of the facing; and ii) large settlements due to active failure in the backfill (England et al., 2000). The test results showing the above are pre-sented later.
A number of bridges comprising a pair of GRS RWs with FHR facing that supports a simple-supported girder via bearings placed on sill beams on the crest of the backfill immediately behind the facing were constructed (i.e., GRS RW bridges; Tatsuoka et al., 1997, 2005). Although this bridge type is more cost-effective than the conventional type, the length of the girder is restricted due to low stiffness of the backfill supporting the sill beams and a low seismic stability of the sill beams because of their small mass. Then, this type was modified by placing a girder on the top of the FHR facings via bearings (Fig. 13a; Tatsuoka et al., 2005, 2009). The first prototype was completed in 2003 for a new bul-let train line in Kyushu (Fig. 15b). The abutment was constructed by the staged construction proce-dure, following the GRS RW technology (Fig. 2). That is, the geosynthetic-reinforced backfill is first constructed, followed by the construction of the fac-ing by casting-in-place fresh concrete. Then a girder was placed on the top of a thin RC facing via bear-ings. In this case, a trapezoidal-shaped zone of back-fill back of the facing was a well-compacted cement-mixed well-graded gravelly soil to minimize the long-term residual deformation and to ensure a very high seismic stability. The conventional type RC
abutment laterally supports the unreinforced backfill and the backfill activates static and dynamic earth pressures on the abutment. In contrast, with this new type abutment, the reinforced backfill laterally sup-ports a thin RC facing, therefore, the backfill does not activate large earth pressure on the facing. After this project, nearly sixty similar bridge abutments were designed or constructed until today. Despite the above, this type of abutment is not free from several problems due to the use of bearings.
a)
b)
Figure 15. a) GRS RW bridge with the girder placed on the
top of the facing via bearings (n.b., the numbers denote the
construction sequence); and b) first prototype at Takada, Kyu-
shu, for a new bullet train line (Tatsuoka et al., 2005, 2009).
3.3 GRS integral bridge
To alleviate all these serious drawbacks with con-
ventional type bridge and also those with integral
bridges and GRS RW bridges, summarized above,
the authors proposed a new bridge type for new con-
struction, called the Geosynthetic-Reinforced Soil
(GRS) integral bridge (Fig. 14a; Tatsuoka et al.,
2008a, b, 2009, 2012a). This bridge type comprises
a girder integrated to a pair of FHR facings (without
using bearings) and backfill reinforced with
geosynthetic layers connected to the facings. This
type of bridge is staged-constructed as follows, re-
ferring to Fig. 16a:
0) When the supporting ground is soft and weak,
the subsoil below the abutments may be im-
proved by, for example, cement-mixing in-place.
1) A pair of GRS walls with the wall face wrapped-
around with geogrid reinforcement is constructed.
2) After major deformation of the supporting
ground and backfill has taken place, thin RC
abutments (i.e., FHR facings) are constructed by
4. Girder
1 1. GRS RW2. FHR facing
3. Movable & fixed
shoes
0. Ground improvement
(when necessary)
casting-in-place fresh concrete on the wall face
wrapped-around with geogrid reinforcement, in
the same way as GRS retaining walls with FHR
facing (Fig. 2a: Tatsuoka et al., 1997).
3) A continuous girder is constructed structurally
integrated to the top of the facings.
Figure 16 GRS integral bridge: a) construction sequences
denoted by the numbers; and b) two-span bridge.
With conventional type bridges and GRS RW
bridges, the length of a single girder is restricted to avoid excessive lateral seismic loads to be activated on the abutment on which a fixed bearing supports the girder. With integral bridges, the girder length is restricted additionally to avoid excessive large cyclic lateral displacements at the top of the facings by sea-sonal thermal deformation of the girder. The girder length of conventional type integral bridge is pres-ently specified to be 50 - 60 m in the USA to restrict the maximum thermal deformation of the girder to about 10 cm. With the GRS integral bridge, such re-strictions as above are looser and its girder length limit would be larger than the value for the conven-tional type integral bridge. Fig. 16b illustrates a GRS integral bridge having two spans.
3.4 NRS integrated bridge
There exist a great number of old conventional type bridges with the girder placed on the abutments via a pair of bearings, that are judged not to satisfy the seismic stability requirement according to the cur-rent design standard. Due to deterioration with time, the number of such bridges as above is increasing every year. It is extremely time-consuming and cost-ly to replace such a bridge with a new one without closing a road or a railway, because the following complicated procedure is necessary: 1) a temporary bridge is constructed adjacent to the existing bridge together with constructing a new approach railway or road; 2) the existing bridge is removed; and 3) a new bridge is constructed at the place of the old bridge; and 4) the temporary bridge is removed.
Figure 17. A new method to reinforce old conventional type
bridges (n. b., the numbers denote the construction sequence).
The authors proposed to reinforce such existing
old conventional type bridges as described above by taking advantage of structural characteristics of GRS integral bridges (Shiranita et al., 2010; Tatsuoka et al., 2012b). This reinforcement method is illustrated in Fig. 17: 1. The girder is initially supported with a steel el-
bow member. 2. The backfill is reinforced with a large-diameter
nail (typically 40 cm in diameter), which com-prises a tie rod (steel or FRP) surrounded with the backfill mixed-in-place with cement-slurry provided from a central hollow rod containing the tie rod (Tateyama et al., 1996). The nails are connected to the top and bottom of the abutment.
3. The girder is integrated to the abutments by plac-ing casting-in-place concrete.
The bridge reinforced by this technology is called the nail-reinforced soil (NRS) integrated bridge. This technology has two major advantages. Firstly, the existing bridge is reinforced while in service, thus this method is much faster and much cheaper than the replacement with a new bridge. Secondly, the NRS integrated bridge is much more stable than the conventional type bridge because of no use of bearings and reinforcement of the backfill.
To validate the advantages of GRS integral bridge and NRS integrated bridge, a series of static cyclic loading tests and shaking table tests were performed on small models in 1g, as described below. In 2009, full-scale models of GRS integral bridge and NRS integrated bridge were constructed. A high construc-tability of these bridge types was confirmed. In the beginning of 2011, full-scale loading tests of these two bridge models were performed.
4. STATIC CYCLIC LOADING TESTS
4.1 GRS integral bridge
Fig. 18 shows the setup of small model tests in the laboratory performed to simulate lateral cyclic dis-
3. GirderStructurally integrated
Firmly connected
1. GRS wall
2. FHR facing
a) 0. Ground improvement
(when necessary)
Gravel
bags
b)
ソイルセメント
セメントミルク
引張芯材
ソイルセメント
セメントミルク
引張芯材Steel or FRP rod
Soil mixed-in-place
with cement
Cement slurry
2. Reinforcing of the backfill
with a large diameter nails
3. Integration of the girder to
the abutment by casting-
in-place concrete
1. Initial support
of the girder
placements at the top of the abutment (or the facing) by seasonal thermal deformation of the girder.
Figure 18. Model tests to evaluate negative effects of lateral
cyclic displacements at the top of the abutment (i.e., facing):
this figure is when the backfill is reinforced (Tatsuoka et al.,
2009).
Figure 19. a) Loading method; and measured time histories
of b) horizontal displacement at the facing top, c) backfill set-
tlement; and d) total earth pressure, unreinforced dense
Toyoura sand (Dr= 90 %) (Tatsuoka et al., 2009). .
Fig. 19 shows the result of a typical test when the backfill is not reinforced and the facing is hinged at the bottom. The top of the facing is displaced cycli-cally on the active side. Despite a fixed relatively small amplitude of displacement, with cyclic load-ing, the earth pressure that increases when the facing is displaced in the passive direction continues in-creasing and the settlement in the backfill that takes place when the facing is displaced in the active di-rection continues increasing, both without showing a sign of stopping. Eventually, active failure takes place in the backfill, while the earth pressure be-comes close to the large passive pressure.
Tatsuoka et al. (2009, 2010b) showed that these trends of behavior are due to the dual ratchet mecha-nism in the backfill (Fig. 20a). That is, only the ac-tive mode of deformation develops when the facing moves in the active direction and it accumulates with cyclic loading, while only the passive mode of deformation develops when the facing moves in the
passive direction and it accumulates with cyclic loading. Fig. 20b illustrates the stress-strain behavior on the active failure plane and passive failure plane during cyclic displacements of the facing. Due to the dual ratchet mechanism, although the facing displac-es with the same amplitude, the shear strain consist-ently increases on both AFP and PFP, which eventu-ally result into active failure and high passive earth pressure.
a)
b)
Figure 20. a) Backfill behind an abutment with vertical
smooth face and a horizontal backfill crest; and b) stress-strain
behaviors on AFP and PFP (Tatsuoka et al., 2010b).
Figure 21. Effects of reinforcing of the backfill with
geosynthetic reinforcement connected to the facing on the set-
tlement in the backfill by lateral cyclic loading of the facing
(Tatsuoka et al., 2009).
Displacement-controlled
loading system
Lateral
displacement
Load cell
Hinge support
129.5
8.5
5 10 20 35
50
.5
Settlement of the crest
Box width: 40
Air-dried Toyoura sand
(Dr = about 90%)
9 separated load cells: LCs
[unit : cm]
8 polyester
reinforcement
layers
Distance (cm) from
the back of the facing
Displacement-controlled
loading system
Lateral
displacement
Load cell
Hinge support
129.5
8.5
5 10 20 35
50
.5
Settlement of the crest
Box width: 40
Air-dried Toyoura sand
(Dr = about 90%)
9 separated load cells: LCs
[unit : cm]
Displacement-controlled
loading system
Lateral
displacement
Load cell
Hinge support
129.5
8.5
5 10 20 35
50
.5
Settlement of the crest
Box width: 40
Air-dried Toyoura sand
(Dr = about 90%)
9 separated load cells: LCs
[unit : cm]
8 polyester
reinforcement
layers
Distance (cm) from
the back of the facing
11.5 cm
L
50.5 cm
L
50.5 cm
0.0
0.1
0.2
0.3
0.4
0.5
0.6
La
tera
l d
isp
lacem
en
t o
f
facin
g, d
/H (
%)
0 5000 10000 150002.0
1.5
1.0
0.5
0.0L= 35 cm
L= 20 cm
L= 10 cm
Se
ttle
me
nt
at th
e c
rest
of
ba
ckfill,
Sg/H
(%
)
Elapsed time, t (sec)
Distance from
the back of facing, L= 5 cm
0 5000 10000 15000
0
1
2
3
4
Initial state
(K0=0.5)
To
tal e
art
h p
ressu
re c
oe
ffic
ien
t, K
Elasped time (second)Elapsed time, t (sec) Elapsed time, t (sec)
Se
ttle
me
nt a
t th
e b
ackfill c
rest,
Sg/H
(%
)
La
tera
l d
isp
lace
me
nt
at th
e fa
cin
g to
p,
d/H
(%
)
Ea
rth
pre
ssu
re c
oe
ffic
ien
t, K
15000
Sg
H= 50.5 cm
a) b)
c) d)
H
Abutment
Lateral displacement, δ
45 / 2o
P
45 / 2o
A
Shear
deformation, us
/ tan mob
Stress ratio,
Peak strength: ( / ) tanpeak
Active direction
3A
1A
2A
0A
5A
4A
7A
6A
3P1P
2P
0P
4P
5P
6P
Passive direction
/ 0 (when K= 1)
Peak strength:
( / ) tanpeak
Relation
on AFP
Relation
on PFP
Passive direction
Active direction
AFP
PFP
R&C: reinforced;
and connection
0 50 100 150 2003.0
2.5
2.0
1.5
1.0
0.5
0.0
-0.5
S5
5 cm
S5
5 cm
Reinforced & No Connected
(R & NoC): D/H= 0.2 %
NR (no reinforcement):
D/H= 0.2 %
Reinforced & Connected (R & C):
D/H= 0.2 %
NR:
D/H= 0.6 %R & NoC:
D/H= 0.6 %
R & C: D/H= 0.6 %
Resid
ual sett
lem
ent
of
the b
ackfill
at
5 c
m
back o
f th
e f
acin
g,
Sg/H
(%
)
Number of loading cycles, N (cycles)
NR: not reinforced R&NoC: reinforced,
but no connection
The result when the backfill is unreinforced
and the facing displacement D/H is 0.6 % pre-
sented in Fig. 21, together with a similar test re-
sult when D/H= 0.2 %, are summarized in Fig. 21.
The results when the backfill is reinforced with
and without connected to the facing are also pre-
sented in Fig. 21. It may be seen that the active
failure in the backfill can be effectively prevented
by reinforcing the backfill with the geogrid layers
connected to the facing. When the geogrid layers
are not connected to the facing, the settlement is
still very large. The passive earth pressure in-
creases even when the backfill is reinforced with
geogrid connected to the facing. Even so, the sta-
bility of the facing is kept, as the facing is sup-
ported by a number of geogrid layers.
4.2 NRS integrated bridge
A series of similar tests as those presented above for the IGS integral bridge were performed on small models of the abutment of NRS integrated bridge (Fig. 22; Tatsuoka et al,, 2012b). To highlight the ef-fects of nailing, the abutment model was embedded only to a depth of 3 m in the supporting ground, simulating prototype abutments not supported with any pile foundation.
Figure 22. Static lateral cyclic loading tests on an abutment
model with nails in the backfill and supporting ground
(Tatsuoka et al., 2012b).
Fig. 23a compares the settlement ratio, S/H, at 5 cm from the abutment, where H is the height of the abutment model (48 cm), plotted against the number of loading cycle, N, when the backfill and support-ing ground was unreinforced or nailed. Fig. 21b compares the outward lateral displacement ratio at the abutment bottom, d/H. Without nailing, distinct active failure took place in the backfill with very large settlements even when D/H was as small as 0.2 %. Besides, due to the development of large pas-sive earth pressure, the abutment bottom was largely pushed out, which restrained the development of high passive earth pressure. These trends became more significant with an increase in the cyclic lateral displacement at the abutment top, D/H. With nails
connected to the top and bottom of the abutment, the active displacement at the abutment footing became essentially zero and the active failure in the backfill was effectively restrained, then the settlement in the backfill became much smaller. Yet, the settlement is not very small due to separated arrangements of nails in the backfill. Some additional measure is necessary to minimize the backfill settlement. Rela-tively high passive pressure developed because the dual ratchet mechanism was still somehow active in spite of nailing. However, the abutment supported by the nails was kept stable.
a)
b)
Figure 23. a) Settlement ratio; and b) lateral displacement ra-
tio at the abutment bottom, plotted against the number of cycle
in the cases with and without nailing (Tatsuoka et al., 2012b).
In summary, the abutment and backfill of a con-ventional type bridge become very stable against cy-clic lateral displacements caused by seasonal ther-mal expansion and contraction of the girder by reinforcing the backfill and supporting ground with nails connected to the top and bottom of the abut-ment. When further the girder is integrated to the abutments, the seismic stability increases substan-tially, as shown below.
Displacement-controlled load
system
Load cell 45
20
180
Units: cm
5 10 20 35
Laser sensor at different distances from the wall
Air-dried Toyoura sand Dr=90 %)
Laser sensor at the top of the wall
Laser sensor at the bottom of the wall
40
Nail reinforcement
Air-dried Toyoura sand (Dr= 90 %)
Displacement transducers Displacement
Transducer
Load cell
Displacement
Transducer
Precise gear
system to control
displacement
1,800
All units (mm)
50 200 (distance from the back of the facing)
100 350
45
02
00
Nail (40 mm in diameter
inclined at 10 degrees
from the horizontal)
Nine local
load cells
0 50 100 150 2004
3
2
1
0
S: measured at the center
between two nails
Nailed backfill
D/H= 0.2%
D/H= 0.4%
D/H= 0.6%
Number of cycles, N
Se
ttle
me
nt
ratio
, S
/H (
%)
(S a
t 5
cm
fro
m t
he a
bu
tme
nt;
& H
= 4
8 c
m)
D/H= 0.2%
D/H= 0.4%
D/H= 0.6%
Unreinforced
backfill
0 50 100 150 200-1
0
1
2
3
4
5
Unreinforced backfill
Range over
Range over
D/H=0.2%
D/H=0.2%
D/H=0.4%
D/H=0.4%
D/H=0.6%
D/H=0.6%
Passive
Active
Late
ral d
isp
lacem
ent
ratio
at
abu
tme
nt
bott
om
, d
/H (
%)
Number of cycles, N
Nailed backfill
Figure 24. Models for shaking table tests (n. b., the whole
model is shown only with GRS-IB) (Muñoz et al., 2012).
5. SHAKING TABLE TESTS
5.1 GRS integral bridge
The following four models were tested (Fig. 22): CB (conventional type bridge): A pair of gravity
type abutments (without a pile foundation) sup-ported the girder via fixed and movable bearings (a hinge and a roller). The backfill was unreinforced air-dried Toyoura sand (Dr= 90 %). Due to a space limitation in the sand box, an additional mass was attached to the girder to simulate a 2 m-long girder. The length scale of the model relative to the as-sumed prototype was 1/10.
GRS-RW (geosynthetic-reinforced soil retaining wall bridge): A pair of sill beams supporting the girder via fixed and movable bearings was placed on the crest of GRS RWs having a FHR facing.
IB (integral bridge): The model was the same as GRS-IB except that the backfill was not reinforced.
GRS-IB (GRS integral bridge): The girder is con-nected to the facings with a pair of L-shaped metal fixtures (20 cm-long, 5 cm-wide and 3 mm-thick),
which did not exhibit the major resisting moment. The backfill was reinforced with grid reinforcement layers connected to the facing. Two other models with a cement-mixed backfill zone immediately be-hind the facing (GRS-IB-C and GRS-IB-C-T) were also prepared. The details are reported in Tatsuoka et al. (2009).
In the first series of shaking tests (on these models described in Fig. 24), twenty sinusoidal waves at a frequency fi of 5 Hz was exerted to the shaking table at each stage increasing the acceleration level incre-mentally by 100 gals (i.e., cm/sec
2) until the models
collapsed. In the second series (on models IB and GRS-IB), input motions at various frequencies were used to examine whether the conclusions from the first series is relevant for a wide range of fi.
Figure 25. Damped SDOF system.
The observed dynamic behaviours of the bridge
models were analyzed by assuming a damped single-degree-of-freedom (SDOF) system (Fig. 25). From the amplification ratio M (= the ratio of the response acceleration at the girder (αt) to the input accelera-tion (αb)) and the phase difference φ between αt and αb observed in every cycle, the tuning ratio β (= the ratio of the input frequency (fi) to the transient natu-ral frequency (f0)) and the transient damping ratio ξ were back-calculated. With each model, the value of β decreased and the value of ξ increased with an in-crease in the number of loading cycle N at the same input acceleration (αb) and with an increase in αb, as typically seen from Figs. 26 - 28. These trends were different among the different models due to their dif-ferent nonlinear load-deformation properties (Munoz et al., 2012). In the first series, all the models started failing when the resonant state was reached. Then, the β value increased exceeding the unity and even-tually collapse took place.
The test results showed that the response accelera-tion for a given input motion decreases and the pos-sibility to reach the resonant state decreases: 1) with an decrease in the initial value of β (i.e., with an in-crease in the initial value of fo, representing the ini-tial stiffness of the structure); and 2) with a decrease
surcharge: 1 kPa
[units:cm]
205.8
60.8
20
Toyoura sand (Dr=90%)
Additional mass (180kg)
Accelerometer
Accelerometer
AbutmentAbutment
Movablesupport
Fixedsupport
Movable
bearingFixed
bearing
Accelerometer
Gravity-type
abutmentLoad cells
CB
[units:cm]
205.8
60.8
surcharge: 1 kPa
Toyoura sand (Dr=90%)
Additional Mass (180kg)
Accelerometer
Accelerometer
35 Reinforcements made of a
phosphor bronze grid
1
3
4
5
6
7
89
2Movablesupport
Fixedsupport
Accelerometer
bearingbearing
GRS RW
60.8
20
Toyoura sand (Dr=90%)
surcharge: 1 kPa
[units:cm]
205.8
Additional mass (180kg)L-shapemetal fixture
Accelerometer
AccelerometerE
mb
ed
de
d a
cce
lero
me
ters
Accelerometer
20
20
43.5
3020105 5
A15 A6 A7 A8
A17
Acceleration extrapolated
A4
A10
A0
Accelerometer
L-shaped
metal
connector
IB
1
2
3
4
5
6
7
8
9
10
60.8
20
Toyoura sand Dr=90%)
surcharge: 1 kPa
[units:cm]
205.8
weight: 200kg
L-shape metal fixture
Model reinforcement made of
a phosphore bronze gridAccelerometer
Accelerometer
35
GRS-IB
AccelerometerL-shaped metal fixtureSurcharge of 1 kPa
Accelerometer Model geogrid
205.8[all units: cm]
85
35
51
Additional
mass: 200 kg60.8
20 35
6
Fixed (x= 0)
Input dis.: ub
Total displacement: ut
Mass: m
Sinusoidal input motion
Inertia = m・d2ut/dt2
Spring: kDashpot: c
Relative dis.: u Acceleration
amplification ratio,
M =
Phase difference, φ
/t b
Tuning ratio:
β=
Measured Unkowns
Damping ratio:
ξ=
= 2πf0
fi=
input freq.(Hz)
f0=
natural freq. (Hz)
in the increasing rate of β (i.e., the decreasing rate of f0) associated with an increase in N and αb. Further-more, 3) for a given input motion, the response ac-celeration decreases with an increase in the damping ratio. Finally, 4) the maximum response acceleration at which the model starts failing increases with an increase in the structural strength. It is shown below that the GRS integral bridge has advantages in all of these factors, therefore, it exhibits a very high seis-mic stability.
Figure 26. Response of model CB (fi= 5 Hz): I, II …. denote
loading stages; and the numerals in ( ) denote the number of
cycle at each stage (Muñoz et al., 2012).
Figure 27. Response of model IB (Muñoz et al., 2012).
Figure 28. Response of model GRS-IB (fi= 5 Hz) (Muñoz et
al., 2012)
a)
b)
Figure 29. Responses of four bridge models (fi= 5 Hz)
(Muñoz et al., 2012)
Fig. 29a shows the relationship between the input
and response accelerations, αt and αb. With CB, M=
αt/αb started increasing at the smallest αb, because of
the largest initial value of β and the largest increas-
0.8
1.0
1.2
1.4
1.6
1.8
2.0
0.0 0.5 1.0 1.5
Resp
on
se r
ati
o o
f accele
rati
on
, M
Tuning ratio, b (b= fi/f0)
I
II
IV
V
VI
VII (1)
VII (20)
1.0
0.8
0.6
0.5
0.4
0.3
x= 0.2
Damping
ratio:
III (2)
III (5)
III (20)
Tuning ratio, b= fi/f0
-2.0
-1.5
-1.0
-0.5
0.0
0.5
0.0 0.5 1.0 1.5
Ph
ase d
iffe
ren
ce,
f(r
ad
)
Tuning ratio, b (b= fi/f0)
0.2
0.3
0.4
0.5
0.6
0.8
1
Series101
Series102
Series103
Series104
Series105
Series106
Series107
I
0.5
0.6
0.8
x= 1.0
Damping
ratio:
0.4
0.30.2
II
III (20)
VI
V
-/4
Tuning ratio, b= fi/f0
Solid curves:
theoretical
First resonant state
Tuning ration, β= fi/fo
Re
sp
on
se
ra
tio
of
acce
lera
tio
n, M
Ph
ase
diffe
ren
ce
, φ
(in
rad
ian
) Loading stage (No. of
cycle at each stage)
-2.0
-1.5
-1.0
-0.5
0.0
0.5
0.0 0.5 1.0 1.5
Ph
ase d
iffe
ren
ce,
f(r
ad
)
Tuning ratio, b (b= fi/f0)
I II III
IV
V (1) VI (20)
VII (1)
VI (1)
V (20)
VII (20)
0.5
0.6
0.8
x= 1.0
Damping
ratio:
0.4
0.30.2
Tuning ratio, b= fi/f0
0.8
1.0
1.2
1.4
1.6
1.8
2.0
0.0 0.5 1.0 1.5
Resp
on
se r
ati
o o
f accele
rati
on
, M
Tuning ratio, b (b= fi/f0)
III III
IV
V (1)
VI (20)
VII (1)
VI (1)
V (20)VII (20)
1.0
0.8
0.6
0.5
0.4
0.3
x= 0.2
Damping
ratio:
Tuning ratio, b= fi/f0
Solid curves: theoretical
Resonant state
Tuning ration, β= fi/fo
Response r
atio o
f
accele
ration,
MP
hase d
iffe
rence, φ
(in r
adia
n)
Loading stage
(No. of cycle at
each stage)
0.8
1.0
1.2
1.4
1.6
1.8
2.0
0.0 0.5 1.0 1.5
Resp
on
se r
ati
o o
f accele
rati
on
, M
Tuning ratio, b (b= fi/f0)
I
IIIVI VIII
VII IX
XXI (1)
XI (7)
1.0
0.8
0.6
0.5
0.4
0.3
x= 0.2
Damping
ratio:
XII
Tuning ratio, b= fi/f0
-2.0
-1.5
-1.0
-0.5
0.0
0.5
0.0 0.5 1.0 1.5
Ph
ase d
iffe
ren
ce,
f(r
ad
)
Tuning ratio, b (b= fi/f0)
I IIIVI
VII
XIX
XI (20)
VIII
XI (1)
XII
0.5
0.6
0.8
x= 1.0
Damping
ratio:
0.4
0.30.2
Tuning ratio, b= fi/f0
Solid curves: theoretical
Resonant state
Tuning ration, β= fi/fo
Response r
atio o
f
accele
ration,
MP
hase d
iffe
rence, φ
(in r
adia
n)
Loading stage
(No. of cycle at
each stage)
0 500 1000 15000
500
1000
1500
GRS IB-T
GRS IB-C
IB
Response a
ccele
ration a
t girder
(gal)
Base acceleration at table (gal)
CB橋桁加速度
GRS RW
GRS IB
CB
1:1
; Moment when failure startes
at resonance
Re
sp
onse a
cce
lera
tio
n a
t th
e g
ird
er,
αt(g
al)
Input acceleration,αb (gal)
Resonance state (start
of failure)
Cement-mixed
backfill part
behind the
facing
0.0 0.5 1.0 1.50
500
1000
1500
CB (stages IV -VII)
GRS-IB-C-T
Response a
ccele
ration a
t th
e g
irder
(gal)
Tuning ratio, bfi/f
0
IB
CB (stages I - III)
GRS-RW
GRS-IB
GRS-IB-C
*
Tuning ration, β= fi/fo
Re
sp
onse a
cce
lera
tio
n a
t th
e g
ird
er,
αt (g
al)
*Response
after failure
*
ing rate of β (i.e., the largest decreasing rate of f0)
with dynamic loading, as seen from Fig. 30. In Figs.
26 - 20, the multiple data points at the same input
acceleration αb with each model denote an increase
in β due to to a decrease in the f0 value with an in-
crease in the number of loading cycles (up to 20).
An increase in β resulted in an increase in M, there-
fore an increase in αt (Fig. 29b). Besides, with CB,
the strength, defined as the response acceleration at
which failure started, was smallest while the damp-
ing ratio at failure was smallest (Fig. 31). Conse-
quently, CB reached the resonant state fastest and
the collapse took place at the smallest αb.
a)
b)
Figure 30. Decrease in fo with an increase in the input accel-
eration of four bridge models (fi= 5 Hz) (Muñoz et al., 2012).
On the other hand, with GRS IB, the initial value
of β was smallest (i.e., the initial value of f0 was
largest) and the increasing rate of β (i.e., the decreas-
ing rate of f0) was smallest (Fig. 30), while the
damping ratio at failure was largest and the strength
was largest (Fig. 31). Consequently, GRS IB
reached the resonant state most slowly and the col-
lapse took place at the largest αb (Fig. 31). This very
high performance of GRS IB is due all to integration
of the girder, the facing and the backfill. The per-
formance of models GRS-RW bridge and integral
bridge (IB) was intermediate due to their intermedi-
ate levels of structural integration. As seen from
Figs. 29 - 31, the dynamic performance of GRS IB
was improved slightly by arranging a cement-mixed
backfill zone immediately behind the facing.
Figure 31. Accelerations and damping ratios at the start of
failure (i.e., at resonance) of six bridge models (fi= 5 Hz)
(Muñoz et al., 2012).
As seen from Fig. 31, the damping ratio, ξ, at the
start of failure increases with an increase in the dy-namic strength and the ξ value of model GRS IB is much higher than model CB. This is because, with an increase in the structural integration among the girder, facing and backfill, the material damping at the start of failure of: i) the material damping of the backfill and supporting ground as part of the bridge system; and ii) the dynamic energy of the girder and abutments dissipated via wave propagation toward the outside of the bridge system increase.
a)
b)
Figure 32. Response of model IB (fi= 2 ~ 20 Hz).
I
IIIII IV
V VI
VII
VIIIIX
X
XIXII
0 200 400 600 800 1000 1200 14000
5
10
15
20
25
30
35
40
Na
tura
l fr
eq
uen
cy o
f b
ridg
e,
f0 (
Hz)
Base acceleration amplitude, Amp[üb] (gal)
GRS-RW GRS-IB
CB
: Resonance state
(fi= 5Hz)
Base acceleration amplitude, Amp[üb] (gal)
0 200 400 600 800 1000 1200 14000
5
10
15
20
25
30
35
40
Base acceleration amplitude, Amp[üb] (gal)
Na
tura
l fr
eq
uen
cy o
f b
ridg
e,
f0 (
Hz)
I
IIIII
IV
V
VI
VII
VIIIIX
X
XIXII
GRS-IB-C
GRS-IB
IB
: Resonance state
(fi= 5Hz)
GRS-IB-C-T
XIII
Input acceleration,αb (gal)
0 500 1000 1500 20000.0
0.2
0.4
0.6
0.8
1.0 ..
..
GRS-IB-T
GRS-IB-C
GRS-IBIB
GRS-RW
damping ratio
Dam
pin
g r
atio x
at
failu
re
Acceleration at resonance (gal)
CB
Response acceleration at resonance, Amp[ut]resonance
(strength)
Base acceleration at resonance, Amp[ub]resonance
A single value of damping ratio
is presented for the respective models.
Response acceleration at resonance, (αt)resonance
(dynamic strength)
Input acceleration at resonance, (αb)resonance
GRS-IB-C-T
0 500 1000 1500
0
500
1000
1500
2 Hz
5 Hz
10 Hz
20 Hz
1 : 2
2 Hz
20 Hz
10 Hz
5 Hz
Resp
onse
Acc
ele
ration , A
r [g
al]
Base Acceleration , Ab [gal]
1 : 1
IB
Input acceleration,αb (gal)
Response a
ccele
ration,
αt(g
al)
10 Hz
20 Hz
2 Hz
Input frequency,
fi= 5 Hz
A Start of failure
0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4
0
500
1000
1500
IB-f
Tuning ratio , b
Resp
onse
Acc
ele
ration ,
t [ga
l]
20 Hz
10 Hz5 Hz
2 Hz
Tuning ration, β= fi/fo
Response a
ccele
ration,
αt(g
al)
Input frequency,
fi= 5 Hz
20 Hz10 Hz2 Hz
A
a)
b)
Figure 33. Response of model GRS IB (fi= 2 ~ 20 Hz).
Figs. 32 and 33 show the results from the second
series on models IB and GRS IB, in which the input acceleration αb was increased changing the input frequency fi at respective stages. In each figure, the data plotted in a broken circle denote the responses at different fi values for the nearly same αb. For the same f0 value, the β value increases with an increase in fi, which results in an increase in the M value be-fore the resonant state is reached and a decrease in M after the resonant state has been passed (i.e., the case denoted by A in Fig. 32). Therefore, with an in-crease in fi, the resonant state is reached at a lower αb and it becomes more likely to pass the resonant state without collapse taking place. It may also be seen from Figs. 32 and 33 that, not only at fi= 5 Hz but al-so at other fi values, the M value increases with an increase in β associated with an increase in the num-ber of loading cycle and αb; and GRS IB was much more dynamically stable than IB. Fig. 34 shows acceleration spectra for a wide range of damping ratio ξ of a typical very severe ground earthquake motion. The aseismic design of RC and steel structures is usually based on a re-sponse spectrum for ξ= 5 %, which is considered representative of the value at failure of these types of super-structure. On the other hand, the ξ value of model GRS IB reached as high as 60 % at the start of failure (Fig. 31). Therefore, the response spectra for ξ up to 70 % are also depicted in Fig. 34. The ini-tial value (before the start of seismic loading) of the
natural period T0 of ordinary prototype bridges, in-cluding ordinary GRS integral bridges, is shorter than the predominant period Tp of severe earthquake motions, which is about 0.35 second in the case of Fig. 34: i.e., the initial value of the natural frequency f0 is higher than the predominant frequency fp=1/Tp (= about 3 Hz). The implications of the shaking table test results are discussed below based on Fig. 34 by assuming that similar response spectrum are kept during a given major earthquake motion.
Figure 34. Acceleration response spectra for different damp-ing rations of NS component of the earthquake motion record-ed at JMA Kobe, 1995 Kobe Earthquake (by the courtesy of Dr Izawa, J., RTRI Japan).
With GRS integral bridges, when compared with
the other bridge types examined in the present study, the initial value of β= fp/f0 (<< 1.0) is lower, which means a lower initial value of the natural period T0 (=1/f0) (<< Tp), which results in a lower initial re-sponse acceleration. Besides, the increasing rate of β is lower, which means a lower increasing rate of T0. Then, the chance to reach the resonant state becomes lower. Furthermore, as the ξ value at failure is very high, even if having reached the resonant state, the response acceleration is kept very low. When ξ= 60 %, the M value at resonance is about 1.35 for the stationary sinusoidal input motions (Figs. 26a – 29a). In the case of Fig. 34, the magnification ratio M at resonance for an irregular input motion defined as the ratio of the response acceleration at T0 = Tp to the value at T0= 0.0 (at which M= 1.0) is also around 1.35. This observation indicates that the design re-sponse acceleration at the girder of a GRS integral bridge can be set to be substantially lower than ordi-nary RC and steel structure (e.g., a bridge girder supported by RC piers). Besides, it is reasonable to set the design maximum response acceleration at the girder of GRS integral brides to be lower than the other bridge types, despite a much higher strength. Based on the above, a design M value between 1.0 and 1.35 is reasonable for such a GRS integral bridge as the one examined in these shaking table tests. Yet, it is necessary to examine whether and how the damping ratio at the start of failure decreas-
0 500 1000 1500
0
500
1000
1500
2 Hz
5 Hz
10 Hz
20 HzResp
onse
Acc
ele
ration , A
r [g
al]
Base Acceleration , Ab [gal]
1 : 2
20 Hz
10 Hz
5 Hz
2 Hz
1 : 1桁での応答加速度
,αt(g
al)
台での入力加速度,α b (gal)
5 Hz
2 Hz
GRS IB
Input acceleration,αb (gal)
Response a
ccele
ration,
αt(g
al)
Input frequency,
fi= 10 Hz Start of failure
20 Hz
0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4
0
500
1000
1500
after failure
GRS-IB-f
Tuning Ratio , b
20 Hz10 Hz5 Hz
Resp
onse
Acc
ele
ration ,
t [ga
l]
2 Hz
5 Hz
20 Hz
2 Hz
Tuning ration, β= fi/fo
Response a
ccele
ration,
αt(g
al)
Input frequency,
fi= 10 Hz
0.1 0.5 1 50
1000
2000
3000
周期 (sec)応答加速度
(gal)
1995兵庫県南部地震神戸海洋気象台 NS成分
h=5% h=10% h=20% h=30% h=40% h=50% h=60% h=70%
Period T0 = 1/(frequency f0) (second)R
esponse a
ccele
ration (
gal) Damping ratio ξ (%)= 5
10
20
30
40, 50, 60 & 70
es with an increase in the girder weight under other-wise the same conditions.
5.2 NRS integrated bridge
The following five models (Fig. 35) were prepared
by reinforcing the conventional type bridge model
(CB) (Fig. 22a):
a)
b)
c)
Figure 35. Small models of conventional type bridge rein-
forced in different ways for shaking table tests, only visible
zones through windows (denoted by solid lines) presented (all
units are in mm: Tatsuoka et al., 2012b).
CB-L (integrated conventional type bridge): The girder was integrated to a pair of gravity-type abutments of model CB using a pair of L-shaped metal fixtures. The backfill was unreinforced air-dried Toyoura sand (Dr= 90 %).
CB-L-N1 (integrated CB with nailing): The backfill and part of supporting ground of model CB-L was reinforced with two layers of 40 cm-long nails on each side. Two nails of each top layer were con-nected to the top of the abutment and two of each bottom layer to the footing toe of the abutment via pin connectors. The model nails were a hollow cir-cular brass rod with a diameter of 4 cm, modeling a full-scale large-diameter nail with a diameter of 40 cm based on a length scale factor of 10. The sur-face of the model nail was made rough by gluing Toyoura sand particles.
CB-L-N2: The length of the bottom nails (40 cm) of model CB-L-N1 were made longer to 60 cm to in-crease the pull-out strength. This model exhibited the highest dynamic stability among those tested in the present study.
Although two other models similar to CB-L-N2
were tested (Tatsuoka et al., 2012b), the results
are not presented herein due to page limitations.
Two series of tests were performed. In the first se-
ries on all the models described above, twenty si-
nusoidal waves at a frequency fi of 5 Hz was exerted
to the shaking table at each stage increasing the ac-
celeration level incrementally by 100 gals. In the se-
cond series on model CB-L-N2, input motions at
various frequencies were used. The observed dy-
namic behaviours of the models were analyzed by
the theory of the SDOF system (Fig. 25). A typical
analysis is presented in Fig. 36. In this test, the shak-
ing was ceased before reaching the resonant state
due to the capacity of the shaking table.
Figure 36. Response of model CB-L-N2 (fi= 5 Hz).
From the results from series 1 (Figs. 36 – 39), it
may be seen that, by integrating the girder to the abutments and further by nailing the backfill and part of the subsoil, i) the initial β value decreased (i.e., the initial value of the natural frequency f0 in-creased); ii) the increasing rate of β (i.e., the de-creasing rate of f0) decreased; iii) the response accel-eration at resonance, at which failure started; increased; and iv) the damping ratio at the start of failure increased, all contributing to a substantial in-crease in the dynamic stability. It may also be seen that the dynamic stability increased by using longer nails. Tatsuoka et al. (2012b) reported that connect-ing the bottom nails to the foundation toe of the
L-shaped
metal fixture
CB-L Accelero-
meter
608 Accelerometer
CB-L-N1
608Accelerometer
Accelerometer
608Accelerometer
AccelerometerCB-L-N2
-2.0
-1.5
-1.0
-0.5
0.0
0.5
0.0 0.5 1.0 1.5
Ph
ase d
iffe
ren
ce, f
(rad
)
Tuning ratio, b = fi/f0
0.2
0.3
0.4
0.5
0.6
0.8
1
Series101
Series102
Series103
Series104
Series105
Series106
Series107
Series108
Series109
Series1
I
0.5
0.6
0.8
x= 1.0
Damping
ratio:
0.4
0.30.2
II
VIIV
III V
VII
VIII
IX
X (20)
0.8
1.0
1.2
1.4
1.6
1.8
2.0
0.0 0.5 1.0 1.5
Resp
on
se r
ati
o o
f accele
rati
on
s, M
Tuning ratio, b = fi/f0
III
IVV
VI
IX
1.0
0.8
0.6
0.5
0.4
0.3
x= 0.2
Damping
ratio:
III
VIII
VII
X (20)
Solid curves: theoretical
End of shaking before
reaching the resonant state
Tuning ration, β= fi/fo
Response r
atio o
f
accele
ration,
MP
hase d
iffe
rence, φ
(in r
adia
n)
Loading stage (No. of
cycle at each stage)
abutment (as models CB-L-N1 and N2) is essential to maintain the contact of the bottom of the footing with the subsoil, thereby to maintain a high dynamic stability.
a)
b)
Figure 37. Responses of five bridge models (fi= 5 Hz).
Figure 38. Decrease in the natural frequency with an increase
in the input acceleration of four bridge models (fi= 5 Hz).
The behavior of model CB-L-N2 for various input frequencies (series 2) is presented in Fig. 40. It may be seen that the magnification of acceleration chang-es by changes in the tuning ratio, β, in the course of dynamic loading in the same way as series 1. This means that the conclusions from series 1, described above, are relevant also for a wide range of fi. In summary, it can be concluded that the dynamic stability of conventional type bridge can be substan-tially improved by the reinforcing technology de-
picted in Fig. 17 (i.e.., the NRS integrated bridge) to the similar level as GRS integral bridge (Fig. 16).
Figure 39. Accelerations and damping ratios at the start of fail-
ure (i.e., at resonance) of four bridge models (fi= 5 Hz).
a)
b)
Figure 40. Response of modelCB-L-N2 (fi= 2 ~ 20 Hz).
3.3 Dynamically brittle versus ductile behaviour
Tatsuoka et al. (1988) reported the performance of various types of retaining wall (RW) during the 1995 Kobe Earthquake. Based on the aboe and those from a series of shaking table tests, they concluded that the dynamic behaviour of gravity-type RWs with unreinforced backfill is generally dynamically very brittle in the sense that the failure starts at a rel-atively low ground acceleration and full collapse may quickly take place subsequently. On the other hand, the dynamic behaviour of GRS RWs with a FHR facing (Fig. 2) is dynamically ductile in the
0 500 1000 15000
500
1000
1500
CB-L-N2 (resonance not reached)
Re
sp
onse a
cce
lera
tion a
t g
irder,
t (
gal)
Base accleration at table, b (gal)
CB応答加速度
; Moment when failure startes
at resonance
CB-L
CB
CB-L-N1
1:1
0.0 0.5 1.0 1.50
500
1000
1500
CB-L-N2
CB (stages IV - VII)
Response a
ccele
ration a
t girder,
t
(gal)
Tuning ratio, b= fi/f
0
CB応答加速度
CB (stages I - III)
CB-L-N1
CB-L
0 500 10000
10
20
30
Na
tura
l fre
qu
en
cy o
f th
e b
rid
ge
, f 0
(H
z)
CB
CB-LCB-L-N1
CB-L-N2
Before
resonance
Base acceleration at respective stages, b (gal)
Resonance state
XIXVIIIVII
VI
V
IV
III
II
I
0 500 1000 15000.0
0.2
0.4
0.6
0.8
Da
mp
ing
ra
tio
x a
t re
so
na
nce
CB
CB-L
CB-L-N1
CB-L-N2
(before resonance)
Acceleration at resonance, (gal)
Base acceleration at resonance, (b)
resonance
Response acceleration at resonance, (t)
resonance
(strength)
0 500 1000 1500
0
500
1000
1500
Resp
onse
Acc
ele
ration , A
r [g
al]
Base Acceleration , Ab [gal]
2 Hz
5 Hz
10 Hz
20 Hz
5 Hz
20 Hz
10 Hz
2 Hz
1 : 2
1 : 1
2 Hz
CB-L-N2
Input acceleration,αb (gal)
Response a
cce
lera
tion,
αt(g
al) Input frequency,
fi= 10 Hz
Start of
failure
20 Hz
5 Hz
0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4
0
500
1000
1500
CB-L-N2-f
Tuning ratio , b
20 Hz10 Hz5 Hz2 Hz
Resp
onse
Acc
ele
ration ,
t [ga
l]
5 Hz 20 Hz2 Hz
Tuning ration, β= fi/fo
Response a
ccele
ration,
αt(g
al)
Input frequency,
fi= 10 Hz
sense that the failure starts at a relatively high ground acceleration and full collapse may take place only after a relatively large increase in the ground acceleration. Based on the above, they proposed to use a higher design seismic coefficient in the quasi-static limit equilibrium-based stability analysis with gravity type RWs than with GRS RWs under other-wise the same conditions.
However, they did not show the mechanism for brittle and ductile behaviours of RWs. The results from the shaking table tests shown above indicate that: 1) the dynamic behaviour of conventional type
bridge is indeed dynamically brittle while the one of GRS integral and NRS integrated bridges is ductile; and
2) the dynamically brittle or ductile behaviour of soil structures results from: a) a low or high ini-tial value of the natural frequency f0; b) a high or low decreasing rate of f0 in the course of dynam-ic loading; c) a low or high response acceleration that the soil structure can survive; and d) a low or high damping ratio of the structure.
a)
b)
Figure 41. A full-scale model of GRS integral bridge con-
structed at Railway Technical Research Institute: a) overall
structure; and b) the left-side abutment under construction.
6. FULL SCALE MODELS AND PROTOTYPES
6.1 Full-scale models
A full-scale model of GRS integral bridge was con-structed during a period of 2008 – 2009 at Railway Technical Research Institute (Fig. 41). The abut-ments were constructed by excavating the backfill of a pair of full-scale models of GRS-RW with FHR facing constructed in 1988 (Tatsuoka et al., 1997). A high constructability of GRS integral bridge was confirmed. The behaviour during construction was
observed and the long-term behaviour is now being observed.
At the end of 2009, a full-scale model of conven-tional type bridge was constructed and the backfill of both of the 6.05 m-high abutments was reinforced with 40 cm-diameter nails, then, the 13.31 m-long steel girder was integrated to a pair of abutments (Fig. 42; Suga et al., 2011). A high constructability of this bridge reinforcement method was confirmed.
In January and February of 2012, full-scale lateral loading tests of the GRS integral and NRS integrated bridge models were performed. The results will be reported in the near future.
a)
b) Figure 42. Full-scale model of conventional type bridge re-
inforced by nailing of the backfill and integrating the girder to
the abutments; a) overall structure; and b) a picture of the mod-
el immediately after completion (Suga et al., 2011).
6.2 Prototype GRS integral bridge
In 2011, the first prototype GRS integral bridge was constructed at the south end of Hokkaido for a new high-speed train line (Fig. 43). The design maximum train speed is 260 km/h. The estimated construction cost is about a half of the one for a box girder type bridge, which is the most conventional solution in this case (Watanabe, 2011). The bridge is heavily in-strumented to observe the behaviour during con-struction and after opening to service. Several GRS integral bridges for railways, including those recon-structed to restore bridges that collapsed by tsunami during the 2011 Great East Japan Earthquake Disas-ter, are now at the stage of planning and design.
Earth pressure cells Earth pressure cells
Cement-mixed backfill Well-graded gravelly soil
5.5
5 m
14.75 m
27 November 2008
13.3
Steel girder
Abutment
Steel girder
Abutment
Integration
Large diameter nails
2.9
Elevation
Plan
2.9
56
.1
[All units in m]
Large diameter nails
a)
b)
Figure 43. First GRS integral bridge, Kikonai at the south
end of Hokkaido, for a new bullet train line (11.7 m wide): a)
overall structure; and b) abutments under construction, summer
2011 (by the courtesy of Japan Railway Construction,
Transport and Technology Agency).
6.3 2011 Great East Japan Earthquake Disaster
The girders and approach fills behind the abutment of a great number of road and railway bridges (more than 300) were washed away by tsunami of the 2011 Great East Japan Earthquake Disaster (Kosa., 2012), as typically seen from Figs. 44 - 46. The above was due to the fact that a girder supported by bearings has a very low resistance against uplift and lateral forces of tsunami while the unreinforced backfill is very weak against erosion by over-flow of tsunami. Connectors and anchors that had been arranged to prevent dislodging of the girders from the abutments and piers by seismic loads could not prevent the flow away of the girders by tsunami forces. These cases showed that the bearings and unreinforced backfill are two weak points of the conventional type bridge not only for seismic loads but also for tsunami current.
Figure 44. A view from the seaside of Namiitagawa bridge,
between Namiitagawa and Kirikiri stations, Otsuchi-cho,
Yamada Line, Iwate Prefecture, East Japan Railway.
Figure 45. A view from the upstream of Tsuyagawa Bridge,
between Motoyoshi and Rikuzenkoizumi stations, Kesen-numa
line, East Japan Raiway.
a)
b)
Figure 46. a) A view from the seaside; and b) the back of the
right bank abutment of Yonedagawa bridge, between Rikuchu-
noda and Rikuchu-tamagawa stations, Noda village, Iwate Pre-
fecture, North-Rias Line, Sanriku Railway.
Figure 47. Collapse mechanisms of conventional type coastal
dyke by over-flowing tsunami current.
6.4 Proposal of tsunami-resistant bridges and dykes
The resistance against the uplift and lateral forces of tsunami of the girder of GRS integral bridge is sub-stantially higher than the conventional type bridge. This is because the girder of GRS integral bridge is integrated to the facings that are connected to the massive approach fill with geogrid layers, while the approach fill that is reinforced and has full-height rigid facings connected to the geogrid layers is very stable against tsunami current. In this case, the ap-proach fill should also have facings on both sides in parallel of the bridge axis to be stable against tsuna-mi current. Even with GRS integral bridges, neces-
GCM: Ground improvement by cement -mixing
5.04 2.21.0
5.42.21.0
0.7 0.7
[All units in m]
GCM
Road surface
12.0
GCMGL= 5.0
Original
ground
Backfill (uncemented)Backfill (cement-mixed
gravelly soil)
6.1
10.75
EastWest RC slab
0.6
0.6
Unreinforced
backfill
Concrete
panel facing
Drainage ditch
Concrete slab on the crest
Leveling pad
Sheet pile
FoundationFoot protection work
Tetrapod
Recurved parapet
Concrete panel facing
Seaside
S
U
sary provisions should be made for protecting the abutments from scouring in the supporting ground by tsunami current.
Most of conventional type coastal dykes comprise unreinforced backfill with concrete panel facings covering upstream and downstream slopes and crest. When subjected to deep over-flow of tsunami cur-rent caused by the 2011 Great East Japan Earth-quake Disaster, at many places, the coastal dykes fully collapsed by the following mechanisms (Fig. 45) and had totally lost their function when attached by subsequent tsunamis. Firstly, a strong tsunami current flowing down the down-stream slope scored in the ground in front of the toe of the downstream slope (as denoted by letter ‘S’ in Fig. 47), by which the concrete facing on the downstream slope was de-stabilized and washed away; and/or the concrete fac-ings at the crest and the upper part of the down-stream slope were peeled off by associated strong uplift forces (as denoted by the letter ‘U’ in Fig. 47) and washed away. Then, the unreinforced backfill was exposed to tsunami current and eroded quickly from the crest and downstream slope. Subsequently, the drawback of a tsunami damaged the seaside slope by the same mechanism as described above. Finally, the full section disappeared.
Figure 48. GRS coastal dykes as a tsunami barrier that can
survive deep over-flow of tsunami current.
On the other hand, coastal dykes that comprise the
geogrid-reinforced backfill with continuous lightly steel-reinforced concrete facings that are firmly con-nected to the reinforcement, like the GRS RWs illus-trated in Fig. 2, should have a much stronger re-sistance against over-flowing tsunami current. Fig. 48 shows several types that can be proposed at this moment. Yamaguchi et al. (2012) performed a series of small model tests and showed that the stability of these types of GRS coastal dykes (Fig. 48) have a very high resistance against flow away by over-flowing tsunami current, much higher than the ordi-nary embankment with concrete facing (Fig. 47).
Tatsuoka and Tateyama (2012) proposed to con-
struct GRS integral bridges (Fig. 16) and GRS em-
bankments/dykes (Fig. 48) to restore the conven-
tional type bridges of railways and roads that col-
lapsed by tsunamis of the 2011 Great East Japan
Earthquake Disaster. Fig 49 illustrates a proposal for
coastal railways. These technologies can also be
used to newly construct coastal railways and roads
that should survive earthquakes and tsunamis.
Figure 49. Coastal railway that can survive a great tsunami
(Railway Technical Research Institute, Japan, 2011).
7. CONCLUSIONS
Geosynthetic-reinforced soil retaining walls (GRS RWs) having staged-constructed full-height rigid (FHR) facing have been constructed as important permanent RWs for a total length of more than 130 km since 1989 until today in Japan, mainly for rail-ways and also for roads and other types of infra-structure. Its current popular use is due to a high cost-effectiveness with high long-term performance and high seismic stability. This success can be at-tributed to: i) the use of a proper type of reinforce-ment (i.e., geogrids for cohesionless soil and nonwoven/woven geotextile composites for high-water content cohesive soil); ii) construction of a FHR facing by such a staged construction procedure that, after the major deformation of the supporting ground and backfill has taken place, fresh concrete is cast-in-place on the wrapped-around wall face of completed geosynthetic-reinforced backfill so that reinforcement layers are firmly connected to the fac-ing; and iii) taking advantage of the rigidity of the facing in design.
A number of embankments and conventional type RWs that collapsed during recent severe earth-quakes, heavy rains, floods and storms were recon-structed to GRS RWs of this type. It was validated that this technology is also highly cost-effective in reconstructing collapsed old soil structures.
The GRS integral bridge was developed, which comprises an integral bridge and geosynthetic-reinforced backfill with staged-constructed FHR fac-ing. The results from model tests in the laboratory,
Planar reinforcement (e.g., geogrid) Planar reinforcement
(e.g., geogrid)
FHR facing (connected to
reinforcement layers)
Planar reinforcement (e.g., geogrid)
Concrete facing (connected to reinforcement,
not allowing the backfill to flow out from openings
Foot protection to
prevent scouring
GRS integral
bridge
GRS RW with FHR
facing
GRS coastal
dyke
Evacuation shelters
along elevated railway
full-scale model tests and construction of a prototype showed the following superior characteristic features over other types of bridge: a) lower cost for con-struction and maintenance; b) less detrimental ef-fects of seasonal thermal deformation of the girder; and c) a much higher seismic stability. Feature c) is due to: 1) a high initial natural frequency; 2) a low decreasing rate of the natural frequency during dy-namic loading; 3) a high energy dissipation capacity at failure; and 4) a high dynamic strength, all by full integration of the girder, the abutments and the rein-forced backfill.
It is argued that GRS integral bridges (with geogrid-reinforced approach fill having facings on its three sides) and GRS embankments/dykes with facings connected to the reinforcement also have a very high resistance against tsunami. It is proposed to reconstruct bridges and embankments of railways and roads that collapsed by tsunami of the 2011 Great East Japan Earthquake Disaster to these types of structures. It is also proposed to newly construct tsunami-resistant GRS integral bridges and GRS embankments/dykes for coastal railways and roads.
8. ACKNOWLEDGEMENTS
The authors would like to sincerely thank their pre-
vious and current colleagues at the University of
Tokyo, Tokyo University of Science and Railway
Technical Research Institute, Japan, for their great
help to a long-term research performed to develop
the GRS technologies described in this paper.
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