Hysteretic Behavior of Conventionally Reinforced Concrete Coupling Beams in Reinforced Concrete Coupled Shear Wall
International Journal of Concrete Structures and Materials
Hysteretic Behavior of Conventionally Reinforced Concrete Coupling Beams in Reinforced Concrete Coupled Shear Wall
This paper presents the experimental results of four full-scale coupling beams in which only horizontal reinforcements are placed, without diagonal reinforcements, with the aim to develop reinforcement details for coupling beams used in connecting side walls in a wall-slab structural system. Each coupling beam specimen was designed according to the deep-beam design procedure that does not use diagonal reinforcements and that is found in current standards. Two cases for basic deep-beam design specimens were investigated wherein (1) U-type reinforcement was added to prevent sliding shear failure of the joints and (2) horizontal intermediate reinforcements were placed. The coupling beam specimens were fabricated with a shear span-to-depth ratio (aspect ratio) of 1.68 and were connected to walls only by horizontal reinforcements, i.e., without diagonal reinforcement. The experimental results indicate that the strength of the beams was about 1.5 times the designed strength of a strut-and-tie model, which suggests that the model is available for predicting the strength of coupling beams with conventional reinforcement layouts such as horizontal and transverse reinforcement bars. The deformation capacity of these conventionally reinforced concrete coupling beams ranged from 1.48 to 3.47% in accordance with the reinforcement layouts of the beams. Therefore, this study found that the performance of conventionally reinforced concrete coupling beams with an aspect ratio of 1.68 can be controlled through the implementation of reinforcement details that include U-type reinforcement and the anchorage of intermediate horizontal bars.
coupling beam; conventional reinforcement; wall-slab structural system; deep-beam design; deformation capacity
Structural walls can serve as an effective structural system to
resist lateral loads, such as earthquakes or winds, in high-rise
buildings. Coupling beams that connect these walls, which
behave independently at each floor, can improve the building’s
lateral resistance capacity. Eurocode 8 (2004) defines a
coupling beam as a beam that is designed to reduce the bending
moment that acts on the wall by about 25% compared to the
case where the wall behaves independently. Figure 1 shows,
however, that a greatly amplified nonlinear deformation in the
coupling beam is required, even if little deformation occurs in
the walls. That is, the coupling beam should not only have the
ability to resist the bending moment but should also possess a
certain deformation capacity because it plays a role in
inducing the ductile behavior of the wall. In this regard, Eurocode 8
(2004) states that a coupling beam can be utilized as the beam
in the moment-resistant frame only in those cases where it is
dominated by the flexural behavior, that is, when the ratio of
the length (l) to the depth (h) is more than 3.0 or the beam can
resist the shear force by diagonal reinforcements.
According to American Concrete Institute (ACI) code
(2014), although a coupling beam can be designed according
to the usual beam design method where l/h C 4, as shown in
Fig. 2, it can nonetheless be reinforced with diagonal cages in
the case of l/h \ 4. Specifically, a coupling beam with l/h \ 2
anffidffiffiffiffiffiffiffiffihffiffioffiffiffisffie factored shear force (Vu) is greater than 4 k
pfck Acw is required to be designed with diagonal cages at the
center of the span. Where fck is the compressive strength of the
concrete and Acw is the cross-sectional area of the web of the
beam. The ACI code also requires that the nominal shear
strength of the beam should be developed only through
diagonal reinforcements and it should be confined sufficiently
by transverse reinforcements, as shown in Fig. 3a. Or, the
entire beam should be confined by sufficient transverse
reinforcement, as shown in Fig. 3b, so that the diagonal
reinforcements cannot undergo compressive buckling.
The basic concept behind reinforcements for coupling
beams for a special structural wall is to confine the beam
using closely spaced transverse reinforcements to prevent
compressive buckling of the diagonal reinforcements.
Accordingly, the currently used codes (ACI 2014; Eurocode
Korea Concrete Institute (KCI) 2012
) for different
areas of the world stipulate that at least four diagonal
reinforcements should be placed and laterally confined. Such
construction details, however, make actual construction
difficult at field sites. Consequently, the current ACI code
(2014) requires that lateral confinement of coupling beams
should be used instead of diagonal reinforcements, as shown
in Fig. 3b, which were recommended in the ACI Committee
318 2011 edition of the ACI code (ACI 2011). These
construction details, however, still pose difficulties in
construction due to the excessive use of lateral confinement
reinforcements and diagonal reinforcements.
In the current standards
(ACI 2014; Eurocode 8 2004; KCI
, all the shear forces and moments that act on coupling
beams are assumed to be borne only by the diagonal
reinforcements. Therefore, transverse and longitudinal
reinforcements that confine the diagonal reinforcements are not
reflected in the design at all. Consequently, all the
reinforcements that are placed in the longitudinal direction of the beam
are cut off at the interface, without being affixed to the walls. In
other words, plastic behavior of the diagonal reinforcements is
encouraged in order to dominate the behavior of the coupling
beams at the ends of such beams where the moment and shear
forces are greatest. Such construction details may induce
sufficient plastic behavior of the coupling beam but may pose
difficulties in construction, as stated earlier, when the shear
and longitudinal reinforcements that are placed for the lateral
confinement of the diagonal reinforcements are excessive in
quantity. In particular, in the case of a wall-slab type apartment
building with no columns, the walls may be as thin as 200 mm
or 300 mm, making rebar placement especially difficult.
Furthermore, the mandatory placement of diagonal
reinforcements, which pose difficulty in construction, makes it
very difficult to perform rebar work in the actual field, which in
turn may lead to faulty construction.
A performance-based design method
Working Group 2010)
was developed recently for the design
of reinforced concrete (RC) members. Thus, the
development of construction details for coupling beams that can be
selected according to the design requirements also is needed.
Specifically, depending on the shear span-to-depth ratio and
shear stress of the coupling beam, diagonal reinforcements
should be used if high deformation capacity is required. If
not, alternative details must be developed. Because such
walls have high stiffness values, the shearing rate of the
lateral force is also high, but the actual deformation is small.
For example, a wall-slab type of structure with no columns
has numerous long walls that experience little deformation.
Therefore, the amount of deformation that is required for the
coupling beams that connect these walls likewise becomes
relatively small. In other words, high deformation capacity is
not required for the coupling beams; thus, suitable
construction details for coupling beams should be investigated.
In order to develop such reinforcement details for coupling
beams that connect walls with less deformation in a
wallslab structural system, this study sought to design coupling
beams in which only conventional reinforcements are
placed, without diagonal reinforcements, and thus to
examine the hysteretic behavior of the coupling beams based on
the results of cyclic loading tests for proposed reinforcement
2. Review of Previous Studies
The significant damage of conventional RC coupling
beams during the Alaska earthquake that occurred in 1964
showed that coupling beams with orthogonal reinforcements
that consist of traditional longitudinal and transverse
reinforcements are susceptible to severe damage under large
Paulay and Binney (1974)
proposed the idea
of placing two intersecting diagonal reinforcement groups,
confined by closely-spaced transverse reinforcements, for
shear-dominant RC coupling beams. Many researchers
(Barney 1976; Shiu et al. 1978; Tegos and Penelis 1988;
Tassios et al. 1996; Galano and Vignoli 2000; Kwan and
Zhao 2002; Fortney 2005; Naish 2010)
have shown that
diagonal reinforcement groups are effective in improving the
strength, ductility, and energy dissipation capacity of RC
short coupling beams. Diagonal reinforcement groups are
generally recognized as the most effective type of
reinforcement detail for providing ductile behavior of RC
short coupling beams that have a span-to-depth ratio of less
than or equal to 2.0. Placing two groups of diagonal bars in
RC coupling beams that have an aspect ratio of less than 4.0
has been specified since 1995 in the ACI Building Code
(ACI 318-95 1995), which specifies that each group of
diagonal bars shall have the same quantity as the transverse
reinforcements for the columns in the special moment frame
to suppress the buckling of each diagonal-bar group.
However, the placement of transverse bars around the diagonal
reinforcement groups, as specified in ACI 318-05, leads to
significant construction difficulties. In order to overcome
such difficulties, ACI 318-08 includes an alternative detail
option where transverse reinforcement is placed around the
beam’s full section, without directly placing transverse
reinforcement around the diagonal bar groups. However, it is
almost impossible to place diagonal reinforcement groups at
a right angle to all the required transverse bars required for
the diagonal confinement and full-section confinement found
in ACI 318-05 (2005) and ACI 318-08 (2008), respectively
To resolve these construction difficulties, several
alternative construction details for RC short coupling beams have
been considered, including rhombic reinforcement, diagonal
reinforcement without transverse ties, bent-up
reinforcement, double beams, and long and short dowel
(Tegos and Penelis 1988; Tassios et al. 1996;
Galano and Vignoli 2000; Hajyalikhan 2015)
none of these construction details allows for performance
that is equivalent to that of coupling beams strengthened
with bundled diagonal reinforcements according to the
existing details specified in the standards and that
significantly improve constructability. Recently, as the building
design concept has changed to performance-based design,
improving constructability and reducing economic costs
have been accomplished by utilizing coupling beams with
proper reinforcement details that are based on the required
deformation capacity under the design loads rather than
based on the aspect ratio, as in the current standard.
More recently, studies have focused on developing
construction details for seismic performance evaluation and
improvement of coupling beams that are strengthened with
horizontal and vertical reinforcements.
Bre n˜a and Ihtiyar
) investigated the effects of different amounts of
longitudinal and transverse reinforcement on the seismic
behavior of four RC coupling beams and discussed the
strength, deformation components, and response parameters
that are needed to construct backbone curves for conducting
nonlinear analyses of a coupled shear wall system.
proposed a simplistic reinforcement
scheme that consists of two separate cages that are similar to
those used for typical beams in RC special moment frames
to minimize the construction problems that are associated
with diagonal reinforcement groups.
reported that proposed details for RC short coupling beams
can transform shear-dominated, brittle behavior into
flexuredominated, ductile behavior.
Cai et al. (2016)
experimental tests using steel fiber-reinforced concrete
(SFRC) coupling beams with conventional reinforcements
and proposed a simplified model that applies the Mohr–
Coulomb failure criterion to predict the seismic shear
strength of SFRC coupling beams. Their test results indicate
that the inclusion of steel fibers can enhance the seismic
performance of SFRC coupling beams and that their
proposed model provides accuracy and reliability.
Lim et al.
investigated the seismic performance of intermediate
aspect ratio coupling beams using proposed alternatives to
mitigate the construction difficulties associated with
diagonal reinforcement by combining conventional and diagonal
reinforcement construction details.
Nabilah and Koh (2017)
tested four conventional RC coupling beams with aspect
ratios of 2.5 and 3.1 and reported that the shear stiffness of
an intermediate length coupling beam was reduced by 0.1%
of the initial stiffness value upon the yielding of the
In Korea since the 2000s, numerous researchers have
conducted studies to evaluate performance and simplify
construction details of coupling beams. For example,
and Yun (2011)
investigated the seismic performance of
strain-hardening cement-based composite (SHCC) coupling
beams that contained different types of reinforcement. They
found that ductile cement-based composites such as SHCC
are effective in improving the ductility and strength of
sheardominant coupling beams.
Shin et al. (2014)
high-performance fiber-reinforced cement composite
(HPFRCC) coupling beams with an aspect ratio of 3.5. Their
test results showed that HPFRCC greatly contributes to the
reduction in crack damage and shear distortion in slender
Jang et al. (2015)
examined the feasibility
of replacing additional transverse reinforcement that is
required for short coupling beams that contain 1.5%
hookedend steel fiber. Also, the
Korea Land and Housing Institute
, 2014) proposed simplified reinforcement details for
transversely confined diagonal reinforcement in short
coupling beams in coupled shear walls.
3. Deformation Capacity Required for Coupling Beams
The behavioral characteristics of coupling beams are
related to the deformation of the lateral forces of the walls to
which the beam is connected, as shown in Fig. 1. The beam
exhibits the deformation of a double curvature due to the
deformation of the left and right walls. In the figure, the drift
required for the beam is the same as the story/floor drift of
the walls. If the stiffness of the left wall is different from that
of the right wall, the drift of the wall with less stiffness will
be greater than the one that is more stiff. In this case, it is
desirable to consider the drift of the wall that has the large
deformation also as the drift of the beam. The drift required
for the walls can be determined by the stiffness values of all
the walls on the floor if the story drift of the floor is
considered the same due to the diaphragm behavior of the floor.
However, determining the drift required for each wall
separately is preferable in order to evaluate whether the coupled
walls provide sufficient ductility with respect to the largest
drift. That is, the drift required for the coupling beams
cannot be less than that required for the walls.
The ACI code (2007) provides a recommendation for the
application of precast concrete walls, which is not covered in
the design criteria, to an area that experiences strong
earthquakes if the performance can be proven through appropriate
performance tests and analyses. In addition, the ACI presents
guidelines for related tests to evaluate whether the structural
members provide sufficient strength, stiffness, ductility, and
energy dissipation capacity via performance testing. These
guidelines also are quoted in the National Earthquake
Hazards Reduction Program (NEHRP) provisions for
seismic regulations for new buildings
(NEHRP 450-1 2003)
Equation (1), defined in NEHRP 450-1, expresses the
required displacement angle as ductility capacity for a wall.
Here, the ductility capacity is the same as for RC, which
means that the ductility should be retained without a rapid
drop in the yield strength with respect to the drift that is (1)
1.5 times the design displacement or (2) from a minimum of
0.8 up to 2.5% according to the shear span-to-depth ratio,
which is the ratio of the height to the length of the shear
wall. Equation (1) is based on a study by
Seo et al. (1998)
that analyzed the experimental results of RC shear walls.
Hidalgo et al. (2002)
reanalyzed the minimum and
maximum displacement angles of the specimens and
adjusted them to 0.8 and 2.5%, respectively. This adjustment was
needed to show maximum shear wall behavior of more than
2.5% in cases of shear wall behavior.
0:67½hw=lw þ 0:5
where hw is the height of the wall for a prototype structure,
and lw is the length of the entire wall in the direction of the
The aforementioned relationship can be used to calculate
the maximum drift required for wall-slab buildings designed
with special shear walls. In South Korea, a typical wall-slab
type apartment usually has a floor height of 2.7 m to 3.0 m
and a wall thickness of 300 mm. Because the frame behavior
governs when the wall length is short or the aspect ratio is
regards these walls as pier walls and
requires that they be designed as columns in a frame. In this
case, the coupling beams should be in accordance with the
strong column/weak beam design principle. If the beams are
designed as coupling beams in special shear walls, the beams
will be overdesigned, and thus, a failure is likely to occur in
wall. The conditions for pier walls are lw/bw B 6 and hw/
lw C 2.0. For wall-slab type buildings in Korea, the
condition in which the wall is not classified as a pier wall is when
the length of the wall is more than
6 9 300 mm = 1800 mm or hw/lw \ 2.0. The range of the
expected hw/lw is 1.5–2.0, and when it is applied to Eq. (1),
the required drift ranges from 1.51 to 1.84, as shown in
Fig. 4. That is, ductile behavior is required for beams that
are connecting the walls within this range, and thus, the
ability to absorb the drift of the walls is required for the
4. Experiment Plan
4.1 Design of the Specimens
Four full-sized coupling beam specimens were designed
and fabricated according to the deep-beam design in the
current KCI standard (KCI 2012). This design is without
diagonal reinforcements and considers constructability and
excess strength. After the basic design according to the
deepbeam design concept, two cases were considered: (1) U-type
reinforcement was added to prevent sliding shear failure of
the joints and (2) horizontal intermediate reinforcements
were placed. Table 1 presents the specimen data, and Fig. 5
shows the construction details for each specimen.
Basically, all the specimens were designed based on the
strut-and-tie model, as shown in Fig. 6. According to the
current standard (KCI 2012), a member, where ln is less than
four times the depth of the member, or on which the load is
applied within a distance of two times the depth of the
member from the support and a compression strut can be
formed between the load-applied point and the support, can
be designed using the struffiffiffitffi-ffiand-tie model if the applied shear
force is less than ð5 k fck =6Þb d and the ratio of the shear
span to depth, ln/d is less than 2.0. The ln/d of each specimen
used in this study was 908/540 = 1.68, which is less than
The amount of reinforcement for each tie was estimated
using Eqs. (2) and (3), and the width of the strut was
calculated using Eq. (4). The minimum requirements for the
vertical and horizontal reinforcements were examined using
Eqs. (5) and (6).
Ast ¼ u fy
n ¼ u Ast fy
wreq: ¼ u 0:85 bs fck b
R bAssi ðsin ciÞ2
0:0025 b s;
where Ast is the required area of reinforcement; Fu,DB, Fu,BC,
and Fu,DE are the forces acting on members DB, BC, and
DE, respectively; u is the strength reduction factor; fy is the
yield strength of the reinforcement; wreq. is the required
width of the compressive strut; bs is the coefficient for the
equivalent stress block; fck is the compressive strength of
concrete; b is the width of the beam; and Asi is the total area
of the distributed reinforcement at spacing si in the ith
direction of the reinforcement that crosses a strut at an angle
ci to the axis of the strut.
In addition, the spacing of the transverse reinforcements
was decided by using Eq. (7) to satisfy the
lateral-confinement condition, as shown in Fig. 5b, so that the cross-ties in
the vertical and horizontal directions would not exceed 200
mm. Appendix provides the design procedures for each
where d is the effective depth.
For the B-1-H specimen, 3-HD22 (1161.3 mm2) was
placed as flexural reinforcement on the upper and lower parts
in accordance with the previously described design process.
The horizontal reinforcement 8-HD10 and the stirrup
HD13@100 were decided using Eqs. (5), (6), and (7) for the
minimum reinforcements in the vertical/horizontal direction.
The horizontal reinforcement 8-HD10 at the center of the
section is the reinforcement for the lateral confinement of the
beam rather than for the flexural strength, but this
reinforcement was affixed sufficiently to the walls so that it
could contribute to the flexure and shear at the joints. In
addition, headed reinforcement was used instead of a
90-degree hook to improve the lateral-confinement
The B-1-HA specimen was almost the same as the B-1-H
specimen, but its reinforcement for the lateral confinement of
the beam was cut off at the beam-wall interface, without
being affixed to the walls. To enhance the confinement effect
by increasing the number of reinforcements,
2-(HD22 ? HD16) (1171.4 mm2) was placed as the upper
and lower reinforcements of the beam; these reinforcements
were affixed to the walls. Unlike the B-1-H specimen,
horizontal reinforcement 8-HD10 was not affixed to the wall, so
these reinforcements did not contribute to the flexure at the
joints. For the beam stirrup, 2-(HD13 ? HD10) was placed
at 100-mm intervals.
The B-2 and B-2-H specimens were strengthened further
using U-type bars to control any sliding shear at the joints
that may occur using the construction details for the B-1-HA
specimen. The number of U-type bars was calculated using
Eqs. (8) and (9). As for the shape of the anchoring cross-ties,
the B-2 specimen had 135- and 90-degree hooks for each
end, respectively, whereas the B-2-H specimen used headed
Shin et al. (2016)
reported the effectiveness
of confining core concrete using headed cross-ties.
Vn ¼ Avf fy l;
l ¼ 1:4;
where Avf is the area of reinforcement for friction design.
0.2 0.4 0.6 1
4.2 Material Properties
Concrete cylinders, 100 mm in diameter and 200 mm in
height, were fabricated for compressive strength tests
conducted at 28 days. The compressive strength value was
found to be 28.7 MPa. Table 2 provides the tensile test
results for the reinforcements used.
4.3 Test Method and Measurements
Figure 7 shows the test set-up wherein coupling beams
were affixed to the lower part of the wall and horizontal
displacement was induced in the upper part of the walls. Ball
jigs were installed on both sides at the center of the upper
part of the wall to prevent out-of-plane deformation when
horizontal displacement was applied. In addition, a roller
was installed on the upper part to slide the upper part of the
wall in the loading direction without rotating it. Bolts were
used to connect the loading frame to the upper part of the
wall, and an actuator with a 1000-kN capacity was installed
at the central axis of the specimen to apply horizontal force.
Strain gauges were attached to the diagonal, horizontal, and
vertical reinforcements to measure the deformation and
yielding period of the reinforcements.
Figure 8 shows that cyclic loading was applied with load
control up to three cycles, and then by two cycles per the
same displacement from drifts of 0.2 up to 6.0%. The loads
of the specimens were measured by a load cell attached to
the actuator. Figure 9 shows the lateral displacement that
was measured by the linear variable displacement
transducers (LVDTs) at the center of the upper and lower
parts of the wall and at the joints. The rotation of the upper
part of the wall was measured also.
5. Experimental Results
5.1 Cracking and Failure Shape
Figure 10 presents the final failure modes of the
specimens. The cracking process for each specimen is described
The failure of the B-1-H specimen revealed that initial
cracks occurred along the interface at the corner of the joint
between the coupling beam and the upper and lower parts of
the wall at the drift of 0.2% and that cracks at the upper right
part of the coupling beam increased at 1.0%. At the drift of
1.4%, many horizontal reinforcements yielded, and the
width of the cracks increased rapidly as the load increased.
Finally, delamination of the concrete cover at the upper part
of the coupling beam occurred at the drift of 2.2%. Failure
did not occur evenly at the joints of the two walls, but was
concentrated only at the part of the joint that was connected
to the upper wall.
For the B-1-HA specimen, initial horizontal cracks
occurred at the center and at the lower left and upper right
corners of the coupling beam at the positive loading of 0.2%
drift. At the drift of 0.4%, the inclined cracks progressed at
the upper and lower parts and at the central part of the beam.
At the drift of 1.0%, the width of the inclined cracks
increased, and delamination of the concrete cover occurred
at the upper and lower parts of the coupling beam. Finally, at
the drift of 3.0%, delamination of the concrete surface
occurred at the upper and lower parts of the coupling beam.
Then, plastic hinges formed at both ends of the coupling
In the case of the B-2 specimen, initial horizontal cracks
occurred at the left interface between the coupling beam and
the lower parts of the wall at the positive loading of 85.23
kN, and diagonal cracks began to appear as the load
increased. In addition, as the diagonal cracks rapidly
increased at the drift of 0.4%, they were found to have been
distributed throughout the beam. At the drift of 1.0%,
delamination in the concrete surface occurred at the upper
left part of the beam, and the width of the diagonal cracks
increased by 1.4%. The test was terminated at the drift of
1.8%. As diagonal tension failure occurred at the center of
the span, not at the interface with the walls, delamination in
the concrete surface near the cracks occurred, leading to
For the B-2-H specimen, initial cracks occurred at the left
interface of the lower part of the wall and the right interface
of the upper part of the wall in the first cycle of positive
loading at the drift of 0.2%, and diagonal cracks occurred
near the center of the coupling beam at 0.4% drift. Finally,
the specimen underwent brittle failure as partial
delamination of the concrete cover occurred at the center of the
coupling beam at the drift of 2.2%. The U-type
reinforcement that was placed to prevent sliding shear failure of the
joints is thought to have contributed to the flexure of the
joint between the coupling beams and the walls so that shear
failure occurred in the beam.
5.2 Load–Drift Curves
Figure 11 presents the load–drift curves of the specimens.
Table 3 shows the yield strength, maximum strength, failure
strength, and displacements at those states for each
specimen. The yield point was set at 75% of the maximum load,
and the failure point was set to the time when the load was
reduced by 25% (0.75 Pm) after the peak point was reached.
The absence of data in the table indicates that the test was
terminated at a state where the load had not reached the point
of 0.75 Pm beyond the peak point. The B-1-H specimen
exhibited a deformation capacity of about 2% drift and
exceeded the design strength, but significant pinching
appeared in the loading zone. The strength of the B-1-HA
specimen also exceeded the design strength, but a reduction
in strength after the peak was more severe in the B-1-HA
specimen than in the B-1-H specimen. The joints between
the beam and the wall were found to have a higher strength
Fig. 13 Plastic behavior that corresponds to formation of
plastic hinges. a Formation of plastic hinge at
wallbeam joints, b Formation of plastic hinge at beam
value than the designed joints even if they were connected
only by the upper and lower horizontal reinforcements. The
B-2 and B-2-H specimens that had been strengthened with
additional U-type reinforcements showed a sharp reduction
in strength after reaching the maximum force, even though
the plastic hinges moved to the beam rather than to the
beam-wall interface. These two specimens eventually failed
with severe x-shaped diagonal cracks, indicating that they all
dominated by shear failure. All the specimens exhibited
significant pinching phenomena at reloading after the load
reversal, thus showing no effective resistance to shear.
Figure 12 presents the envelope curves of the specimens.
The B-1-H specimen exhibited the greatest strength and
Fig. 14 Comparison of dissipated energy values: a
accumulated energy, b cyclic energy.
ductility capacity, showing a deformation capacity of about
2% drift. The B-2 specimen, to which U-type reinforcements
were added, showed lowest strength and deformation
capacity. As shown in Fig. 13, if a plastic hinge forms at the
wall-beam joint, the deformation will be dispersed, but if it is
concentrated at the center of the beam, brittle failure is likely
to occur because the shear deformation is concentrated. In
the case of the B-2 and B-2-H specimens with U-type
reinforcements at the joints, plastic hinges formed at the
center of the beam, because the wall-beam joints were
strengthened. Therefore, the overall behavior seemed to be
extremely brittle as the deformation was concentrated at the
center of the beam in the form of shear deformation.
The strength of a coupling beam that is connected to walls
with only horizontal reinforcements, without the use of
diagonal reinforcements, was calculated based on the
strutand-tie model. As shown in Table 3, this strength value was
found to be 1.21–1.64 times higher than the design strength.
Except for the B-2 specimen with a significantly lower
strength value in the negative direction, the ratios were
1.46–1.64 times higher than the design strength, indicating
that the average is about 1.5 times the design strength. The
drift of the specimens at the maximum load ranged from
1.13 to 2.11%. Compared to the B-1 series specimens whose
horizontal reinforcements of the beams were anchored in the
walls, the B2 series specimens whose main bars were
anchored in the walls showed low drift percentages. The
wall drift of Eq. (2) becomes 1.63% when it is considered as
the required drift until failure. When this value is compared
with the test results, all the specimens, except for the B-2
specimens, can be said to have exceeded the required drift of
1.63%. Of course, more than 1.5 times the design response
displacement of the building is required for an actual
building, which needs to be taken into consideration.
5.3 Energy Dissipation Capacity
Figure 14 presents the dissipation energy and accumulated
energy of the specimens for each displacement step. On the
whole, the cumulative dissipation energy values of all the
specimens are very similar. The B-1-HA specimen shows
similar dissipation energy up to the drift of 2% despite
slightly low strength, as the number of horizontal
reinforcements anchored in the wall is less than that of the
B-1H specimen. In the case of the B-2 specimen, the energy
value per cycle is lower than for the other specimens. This
outcome is probably due to its brittle behavior whereby the
strength suddenly decreased after the maximum load was
reached as shear failure occurred in the beam.
5.4 Reinforcement Strains
Strain distributions of the horizontal reinforcements were
observed at the wall-beam joints. The specimens with
horizontal reinforcements anchored in the wall exhibited a
similar degree of deformation as the interface with the wall
and at the vicinity of 0.5 d (d: effective depth of coupling
beam). Figure 15 shows the deformation of the horizontal
reinforcements located at 0.5 d from the wall-beam joints.
The B-1-H specimen exhibits low deformation of the
horizontal reinforcement at the center, but the B-2 and B-2-H
specimens with U-type reinforcements show greater
deformation of the reinforcements at about 1% drift. This
outcome is due to the fact that the flexural deformation caused
by the upper and lower horizontal reinforcements was
dominant at the wall-beam joints in the case of the B-1-H
specimen, whereas the shear deformation at the center of the
beam dominated for the B-2 and B-2-H specimens. This
phenomenon also can be seen in Fig. 16 that shows the
stirrup strain response of each specimen. For the B-2 and
B2-H specimens, the stirrup at the center yielded, and the
deformation increased sharply after the drift of 1%. In the
case of the B-1-H specimen, the strain of the stirrup at the
center of the beam increased, but the degree of increase was
less than that of the specimens with U-type reinforcements.
Stirrup of beam
Location from wall face (mm)
Location from wall face (mm)
Stirrup of beam
Stirrup of beam
Location from wall face (mm)
Location from wall face (mm)
For the B-1-HA specimen, the overall stirrup strain was low,
and was remarkably low especially at the center of the beam.
This finding suggests that in cases where the wall and the
beams are connected only by upper and lower horizontal
reinforcements, the behavior is dominated by the flexural
behavior of the joints, and the stress from the wall may not
be transferred properly to the coupling beam when the
connection is weak.
From the comparison of the B-2 and B-2-H specimens
tested to investigate the confinement effect of headed cross
ties, the cross ties with the standard hook shows a slightly
greater strain than the headed cross ties. Therefore, no
further increase in the confinement effect is likely when headed
bars are used as the horizontal and vertical cross ties.
An experiment on four full-scale coupling beam
specimens was conducted to investigate the hysteresis
characteristics of the coupling beams with horizontal reinforcements
instead of diagonal reinforcements for wall-slab structural
system. The study used reinforcement details for the
connection between the beam and the wall as variables, based
on reinforcement details required by current standards, to
investigate the behavioral characteristics of coupling beams
whose shear span-to-depth ratio is less than 2.0 and where
only conventional reinforcement layouts are placed, i.e.,
without diagonal reinforcements. The following conclusions
were drawn based on the test results.
Coupling beams whose shear span-to-depth ratio was
1.68 and which were connected to walls only by
horizontal reinforcements, without diagonal
reinforcement, showed strength that is about 1.5 times the
design strength for a strut-and-tie model, thus
indicating that proper design strength is possible using these
construction details. Overall, the deformation capacity
was about 2%, which indicates a certain amount of
deformation capacity. A pinching phenomenon,
however, that occurred after the load reversals indicated a
low level of energy dissipation.
Horizontal reinforcements that were anchored in the
walls for the lateral confinement of the beam led to an
increase in beam strength. Even in cases where part of
the horizontal reinforcement was not anchored into the
wall, the design strength and a certain degree of
ductility capacity were provided, and plastic hinges
could be induced completely in the wall-beam joints.
However, the strength gradually decreased after
reaching the maximum force, which suggests that if the wall
and beams are connected only with upper and lower
horizontal reinforcements, the overall behavior is
dominated by the flexural behavior of the joints, and
the stress from the wall may not be transferred properly
to the coupling beam when the connection is weak.
A comparison of the B-2 and B-2-H specimens
indicates that the connection reinforcement used in
standard hook construction details can lead to a slightly
high strain distribution. Therefore, no further increase
in the confinement effect would be expected in cases
where headed reinforcements are used as the
When U-type reinforcements were placed at the joints
to control slippage due to the plasticization of the
joints, excessive shear deformation occurred as plastic
hinges were induced into the center of the beam,
without forming at both ends. Consequently, brittle
failure occurred when only horizontal reinforcements
This research was supported by a grant from the Korea Land
& Housing Institute. And this work was also supported by
the Brain Korea 21 Plus Project of Dept, of Architectural
Engineering, Chungnam National University in 2017.
Appendix: Design of specimen
ln ¼ 908 mm; B
D ¼ 300 mm
fck ¼ 24 MPa; fy ¼ 400 MPa;
ln=D ¼ 908=590 ¼ 1:54\2:0 ð)Deep beamÞ;
Vu ¼ 408 kN
1. Design loads
b; T ¼ Az þ ff
As fy ð1174Þ
a ¼ 0:85 fck bw ¼ 0:85 24
¼ 76:73 mm from C ¼ T
Member force in strut-tie model as shown in Fig. 6
RME ¼ ð408 103 454Þ
¼ 0; )FAB ¼ 400:59 kN
RV ¼ ð408
¼ 0; )FBD ¼
6. Strengths of strut and node in beam
wreq ¼ u
wreq;BD ¼ 0:75 0:85 0:8 24 300 ¼ 155:7 mm
160 mm ðNode D, bs¼ 0:8; CCT)
wreq;CE ¼ 0:75 0:85 0:6 24 300 ¼ 207:62 mm
210 mm ðNode E, bs¼ 0:6; CCT)
wreq;Tie ¼ 63:076 mm
x ¼ qffiðffiwffiffiffiffirffieffiqffiffi;ffiTffiffiiffieffiÞffiffiffiffiffiþffiffiffiffiffiðffiffiwffiffiffirffieffiffiqffi;ffiDffiffiffiBffiffiÞffi
h ¼ tan 1
ð63:076Þ2 þ ð155:7Þ2 ¼ 167:9 mm
7. Check for the minimum reinforcements in beam
8. Calculation of maximum strength (B-1-HA specimen)
When concrete reaches its ultimate strength due to
Mn ¼ As
¼ 229:09 kN m
For specimens B-1-H, B-2 and B-2-H that all of
the horizontal bars of beam (6-HD22 ? 8-HD10)
are anchored in wall, Mn is achieved from a sectional
analysis (using Structural Analysis Program XTRACT)
as shown in below. The forces of struts
and ties are can be calculated by same process by using
When stirrup (2-(HD13?HD10)@100) reaches its
yield strength due to shear force,
Therefore, it is expected that the behavior of the
specimen will be governed by bending moment.
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