Experimental Cyclic Behavior of Precast Hybrid Beam-Column Connections with Welded Components
International Journal of Concrete Structures and Materials
Experimental Cyclic Behavior of Precast Hybrid Beam-Column Connections with Welded Components
Sadik Can Girgin 0
Ibrahim Serkan Misir 0
Serap Kahraman 0
0 Department of Civil Engineering, Dokuz Eylul University , 35160 Buca-Izmir , Turkey
Post-earthquake observations revealed that seismic performance of beam-column connections in precast concrete structures affect the overall response extensively. Seismic design of precast reinforced concrete structures requires improved beamcolumn connections to transfer reversed load effects between structural elements. In Turkey, hybrid beam-column connections with welded components have been applied extensively in precast concrete industry for decades. Beam bottom longitudinal rebars are welded to beam end plates while top longitudinal rebars are placed to designated gaps in joint panels before casting of topping concrete in this type of connections. The paper presents the major findings of an experimental test programme including one monolithic and five precast hybrid half scale specimens representing interior beam-column connections of a moment frame of high ductility level. The required welding area between beam bottom longitudinal rebars and beam-end plates were calculated based on welding coefficients considered as a test parameter. It is observed that the maximum strain developed in the beam bottom flexural reinforcement plays an important role in the overall behavior of the connections. Two additional specimens which include unbonded lengths on the longitudinal rebars to reduce that strain demands were also tested. Strength, stiffness and energy dissipation characteristics of test specimens were investigated with respect to test variables. Seismic performances of test specimens were evaluated by obtaining damage indices.
beam-column connections; precast concrete; welding; unbonded length; damage index
One-story precast concrete structures constitute a
significant part of industrial buildings in earthquake-prone regions
in Turkey. Post-earthquake observations revealed that
beamcolumn connections are widely influence the overall seismic
response of precast concrete structures (Saatcioglu et al.
2001; Ozden and Meydanli 2003; Senel and Palanci 2013).
It is still a challenging subject to develop precast concrete
beam-column connections emulating the seismic
performance of monolithic systems to maintain advantages of
precast construction process for multi-story buildings.
In the literature, joint is defined as the intersection of beam
and column elements while connection is the region where
precast elements connected with a technique (welding, bolting
etc.) during construction. Moment-resisting beam-column
connections can be categorized mainly as emulative (wet) and
dry connections (e.g. ACI 550.2R 2013). In addition,
emulative and mechanical components can be assembled to
constitute hybrid beam-column connections (Negro and Toniolo
2012) which are commonly used in Turkey.
In emulative connections, continuity of reinforcing bars is
provided by a coupling connector or splicing throughout the
designated gaps in precast beam and column elements at a
precast construction facility. Precast beams are firstly
supported on columns’ cover concrete and then the topping
concrete is poured on site to fill the gaps in the column and
the top of the beam (Park and Bull 1986; Chen et al. 2012).
Im et al. (2013) tested five interior precast beam- column
connections with U-shaped beam shells. Main test
parameters were seating length of beam to joint, steel angle for
cover concrete and installation of headed rebars. It was
concluded that increase in effective depth of beam-column
connection can be obtained by decreasing the seating length
and the beam shell thickness.
Dry beam-column connections are achieved by connecting
the precast elements with post-tensioning, welding or
embedded rods. Chang et al. (2013) presented experimental
results of two full-scale interior beam-column connections
with embedded ductile rods within the joint region. Main test
parameters were the use of high strength concrete,
posttensioning and high performance reinforcing steel. They
were concluded that specimens sustained large drifts without
any strength degradation. Experimental studies on
post-tensioned connections showed that increase in mild
reinforcement ratio results in improved ductility and energy
dissipation capacities. However, in some cases premature
buckling and rupture of mild reinforcement may occur
before reaching the drift ratios that the connection can
sustain (Cheok et al. 1993; Priestley et al. 1999; ACI T1.2
2003; Ertas et al. 2006). It is proposed that, unbonding of
reinforcing bars over a length of precast beam and column
elements in a sleeve can reduce the strain demands and
prevent rupture which leads significant degradation in
strength (Cheok et al. 1996; Pampanin et al. 2001; Belleri
et al. 2012). However, unbonded length should be selected
properly to ensure yielding of reinforcing bars without
premature rupture (Cheok et al. 1993).
In the third category, hybrid (emulative-welded)
connections are widely used in residential and industrial frame
structures where negative moment continuity is made
through cast-in-place connection while positive moment
continuity is satisfied through welding. Figure 1 shows a
fabrication process of a precast structure by applying hybrid
connections with the stages from precast facility to site.
First, beam longitudinal and transverse rebars are welded to
beam end plates and the remaining rebars in beam elements
are installed (Fig. 1a, b). At the meantime, precast columns
with corbels of hybrid connections are prepared and a gap is
deliberately left at the top of the joint before casting of
concrete to fill with topping concrete (Fig. 1c). After
transportation and installation of precast beam, column, and slab
elements, continuity between these elements are provided by
pouring topping concrete in each floor level (Fig. 1d).
Seismic behavior of precast connections is required to be
proved by experimental or numerical studies in terms of
equivalent strength and ductility reflecting the monolithic
behavior as specified in seismic codes (TEC 2007; ACI 318
2011). There has been a great deal of research by means of
experimental and numerical studies on reinforced concrete
beam-column joints (Ronagh and Baji 2014; Kim and
Hyunhoon 2015; Kassem 2015; Rashidian et al. 2016; Lim
et al. 2016). However, there are a limited number of
experimental studies on the behavior of hybrid
(emulativewelded) connections in the literature. Ertas et al. (2006)
tested half-scaled one exterior beam-column connection
where the first diagonal crack near the connection was
reported at 2.2% drift ratio and the beam bottom rebars
ruptured at the first cycle of 3.5% drift ratio. Yuksel et al.
(2015) tested five half-scaled exterior hybrid connections
including a slab which are subjected to monotonic and cyclic
drifts applied at the beam tip. Monotonic and cyclic tests on
first three specimens showed that strength degradation
occurs due to rupture of welded longitudinal rebars and
transverse rebars at the connection. Improved specimens
showed an increased energy dissipation capacity while the
observed in-cycle degradation was about 50% at a drift ratio
of 3%. During the tests, increase in strain demands of beam
welded rebars played an important role for overall behaviors
of the connections.
Experimental studies on welded ASTM A615 type
reinforcing bars (ASTM 1992) revealed that welding process
causes heat-affected zone on reinforcements following an
embrittlement on the material which is undesirable for a
ductile seismic design (Rodriguez and Rodriguez 2006;
Rodr´ıguez and Torres-Matos 2013). This kind of rebars
showed a decreased tensile strain capacity in the vicinity of
the welded region. However, welding is still a common used
technique in connecting the steel parts of structural elements
Fig. 1 A precast structure under construction with hybrid beam-column connections. a Welding of longitudinal beam bottom rebars
to beam end plate. b Installation of flexural and shear reinforcement for a precast beam. c Installation of flexural and shear
reinforcements to constitute the precast column and corbels. d Assembling precast beam and column elements on site.
in precast industry. As shown in the latter sections in this
study that, rebars with low carbon content (e.g. B420C grade
steel) defined in TS708 code (2010) could show a sufficient
ductile behavior under tensile forces after a welding process.
An experimental research program was carried out to
improve the cyclic behavior of hybrid (emulative-welded)
beam-column connections. In this study, half scale one
monolithic and five precast specimens representing interior
beam-column connections were tested under reversed cyclic
loading. Strength, stiffness and energy dissipation capacities
of test specimens were investigated with respect to welding
coefficient and unbonded length as the main test variables.
Moreover, damage indices were also obtained to compare
the seismic performance of these test specimens.
2. Experimental Study
2.1 Material Tests
Weldability and hence the mechanical properties of a
reinforcement after welding depend on the chemical
composition of the material defined as carbon content (C) and
carbon equivalent (CE) ratios (Atakoy 2014; TS 708 2010;
ASTM A706M 2013). Therefore, appropriate type of
longitudinal rebars which is compatible with the upper limits
specified by the TS 708 code (2010) (C: 0.22%, CE: 0.50%)
were installed to beam and column sections except one
specimen. The rebars were also satisfying the carbon content
and CE allowable limits (C: 0.30%, CE: 0.55%) given in
ASTM A706M (2013) for Grade 60 rebars which has similar
In order to characterize the tensile behavior of the rebars
welded to steel plate representing the rebar-plate connection
subassembly in the test specimens, tensile test was
performed on 18 mm diameter reinforcing bars welded from
both sides to PL plate (St 37 steel) as shown in Fig. 2a.
Appropriate electrodes (E42 type) were used to weld rebars
to plate providing filler metal requirements due to TS EN
ISO 2560 (2013). Figure 2b shows the specimen during
testing while average strains in rebar were determined by
using optical sensors focused on two points located at the
vicinity of the plate.
Rebar rupture occurred within the unwelded part of the
rebar specimen at 9.5% strain with a clear necked region
verifying a ductile behavior under tension as shown in
Fig. 2c. A tensile test was also performed on single rebar
(not welded to plate) where tensile strain at rupture is
determined as 12%. While interpreting the decrease in
tension strain in the case of welded rebar, potential stress
concentrations near the plate and heat-affected zones should
be keep in mind. Besides, it is concluded that using proper
type of reinforcement didn’t lead to a brittle failure. The
stress–strain curves gathered from the tensile tests for single
and welded rebars are shown in Fig. 3.
Tensile test results of longitudinal and transverse rebars
are given in Table 1. Average cylindrical concrete
compressive strength (fcm) for the monolithic specimen, for beam
and column precast elements, and the topping concrete were
37, 42, and 35 MPa at the test days.
2.2 Test Specimens
Half scale test specimens were designed in accordance with
Turkish Earthquake Code (TEC 2007) and ACI 352R (2002)
requirements representing an interior connection between
inflection points of beam and column elements of a four-story
precast building with moment frames of high ductility level.
For interior beam-column joints, bond requirements are
specified by ACI 318 (2011) (Eq. 1) and ACI 352R (2002)
(Eq. 2) for column depth-to-beam rebar diameter ratio (hc/db)
which are given in the following equations.
Fig. 3 Tensile test results for an 18 mm diameter single rebar
and a rebar welded to plate.
For hc (= 400 mm), db (= 18 mm) and fy (= 461 MPa)
monolithic beam-column joint fulfils the bond requirements
in Eqs. (1) and (2).
Applied axial load was 8% of axial load capacity of the
column during cyclic reversals. In order to investigate the
failure modes in the vicinity of the connections, columns
were designed to have a restricted damage during testing.
Ultimate moment ratio of column-to-beam was 2.4.
Longitudinal rebars were 22 mm in diameter for columns and
18 mm for beams, while 10 mm rebars were used for
transverse reinforcement Geometry and reinforcing details of
the monolithic specimen are shown in Fig. 4.
Geometry, reinforcing and welding details of the precast
connections are also shown in Fig. 5. During the first stage
of the study, SP1, SP2, and SP3 specimens with various
welding coefficients (a) were tested. For welded connections
of precast concrete structures, seismic connection forces are
multiplied by a welding coefficient (a) to obtain the design
forces in the scope of Turkish Earthquake Code (2007).
Welding coefficient was recommended as 2 and 1.5 in TEC
(1998) and TEC (2007) codes, respectively. Therefore, the
required area for welding between beam bottom longitudinal
rebars and beam support plate (PL-1) shown in Fig. 6a was
calculated by a 9 Fy where Fy is the yield force of a beam
bottom longitudinal rebar. Corbel lengths (Lc) of specimens
were determined considering the welding length (Lw) of
beam longitudinal rebars as given in Table 2. Minimum
welding thickness (amin) was considered as 3 mm. Two
additional vertical plates (2 9 PL-2) were welded to beam
end plate (PL-1) to provide an extra welding area for stirrups
in order to improve the connection between the stirrups and
the end plate as shown in Fig. 6b.
Before the construction process of the test specimens at
the laboratory shown in Fig. 7, beam and column elements
of precast specimens were constructed separately in a precast
facility and then transported to the laboratory. Columns were
kept in a vertical position and Beam 1 and Beam 2 elements
were supported on the corbel plates (PL-3). Beam end plates
(PL-1) were welded to corbel plate (PL-3) at both sides with
a 7 mm thick welding. 20 mm gaps were deliberately left
between beam and column vertical faces at each side of the
column and filled with grout properly (Fig. 7a). Beam top
longitudinal rebars were placed and a topping concrete with
a thickness of 150 mm was casted for the completion of
specimens (Fig. 7b).
Geometry and reinforcing details of the precast specimens
are given in Fig. 8. Free ends of the beams were fixed in
vertical direction, 1800 mm far from the column axis while
column was supported with a hinge at the bottom. Shear
span to beam depth (a/d) ratio for the specimens are about 3.
At the second stage, SP1 and SP3 specimens were
improved considering the failure modes observed in the
previous tests. SP1-R and SP3-R represents the revised
specimens by adapting the unbonded length concept,
generally applied in post-tensioned connections. Unbonded
lengths (Lu) for precast connections are proposed between
100 and 200 mm or 4db–8db, where db is longitudinal rebar
diameter by ACI 550.2R (2013) as a design guide for precast
jointed systems. Unbonded lengths in the revised specimens
were considered as 5db and 10db for SP1-R and SP3-R
specimens, respectively. Turkish Earthquake Code (2007)
specifies transverse reinforcement spacing as a minimum of
8db, 150 mm and 0.25hb, where hb is the beam depth. For
SP3-R specimen, stirrup spacing (sh) of beams was
decreased to 75 mm and additional ties were installed to
reduce possible buckling of rebars. Application of unbonded
length and reinforcing details of SP3-R specimen which has
a decreased welding coefficient but an increased ratio of
additional ties compare to SP1-R is shown in Fig. 9.
Table 2 summarizes the welding coefficient (a), corbel
length (Lc), transverse reinforcement ratio (qw), slenderness
ratio of longitudinal rebars (sh/db), applied unbonded length
(Lu), the carbon content (C) and the CE ratio of the
longitudinal bars corresponding to each precast specimen.
Table 1 Tensile test results for longitudinal and transverse rebars.
Fig. 4 Geometry and reinforcing details for monolithic specimen a beam section, b column section, and c interior connection.
2.3 Experimental Set-up
Location of a specimen in the test set-up is shown in
Fig. 10. Experimental set-up was designed and the tests
were performed based on the criteria stipulated by ACI
374.1 (2005). Drift ratios are defined considering the
uncertainties in strong ground motions and structural
properties for beam-column connection sub-systems. Therefore,
tests were continued up to 3.5% drift ratio to observe the full
post-yield behavior of the connections required by ACI
Lateral cyclic displacement protocol adopted based on
ACI 374.1 (2005) were applied to top of the column with an
increasing amplitude from 0.15 to 5.0% drift ratios. Drift
ratio is defined as D/Hc, where D is lateral displacement and
Hc is the height of the column between hinges. Predefined
loading protocol is shown in Fig. 11.
Fig. 7 a Installation process of precast beam-column connection specimens and b casting of topping concrete.
Fig. 8 Geometry and reinforcing details of precast specimens, a beam cross section, b column cross section, and c corbel details.
In order to monitor deformations, 12 linear variable
displacement transducers (LVDT) were installed by 300 mm
spacing in vertical and horizontal directions. Installation of
LVDT’s are shown in Fig. 12a for MONO, SP1-R, SP3; and
in Fig. 12b for SP1, SP2, SP3-R specimens. Also, two string
pots were installed at the top of the column to monitor the
Fig. 9 a Application of unbounded length and b beam cross-section details of SP3-R specimen.
Fig. 10 Components in experimental set-up: (1) axial load
frame, (2) hydraulic jack, (3) hinge, (4) actuator, (5)
load cell, (6) reference frame, (7) pendulum support,
(8) load cell, (9) fixed support, (10) out-of-plane
support frame, (11) steel reaction frame.
Fig. 11 Loading protocol.
applied top displacements. Strain gauges were mounted on
beam and column longitudinal and transverse rebars to
monitor strains throughout the tests as shown in Fig. 13.
3. Experimental Results
3.1 Damage Patterns
Tests were continued up to certain drift ratios where the
connection specimens showed significant strength
degradation during loading. Figure 14 shows the damage patterns at
each drift ratio for monolithic and precast connection
specimens. In Fig. 14, concrete crushing is represented by
shaded areas. Flexure, shear and bond failures were
concentrated near the joint of the MONO specimen as shown
in Fig. 14a. In the MONO specimen, bond slip of beam
longitudinal rebars occurred after 2.75% drift ratio.
For precast connection specimens, flexure and shear
cracks occurred along the beams while columns had only
hairline joint cracks. During the experimental studies, rebar
buckling and the subsequent rupture of welded beam
longitudinal rebars were observed by mainly a visual inspection
with the spalling of cover concrete. Damaged specimen
photos showing the initiation of rebar buckling and the
rupture of rebars for SP1 specimen is given in Fig. 15. Drift
ratios corresponding to initiation of each damage state
observed on the specimens throughout the experimental
study are given in Table 3.
Beam longitudinal rebars of precast specimens welded to
end plates buckled and sequentially or subsequently ruptured
in most cases except for improved SP3-R specimen. Since
SP1 specimen had welded rebars with CE similar to ASTM
A615 type rebars (ASTM 1992) as shown in Table 2, rebar
rupture occurred at a relatively lower drift ratio (2.2%)
compared to revised precast specimen (SP1-R). SP2 and SP3
specimens had lower carbon content and concrete crushing
at 2.75% drift (Fig. 14d) followed by rebar buckling and
rupture were observed in these specimens In SP1-R
specimen, provided additional ties were improved the behavior
and prevented early rebar buckling. Moreover, the
combination of applying unbonded length and additional vertical
ties could postpone the rebar buckling in SP1-R and SP3-R
specimens, and subsequently the rebar rupture.
3.2 Lateral Load–Drift Ratio Relationships
Lateral load–drift ratio relationships for monolithic and
precast connection specimens are shown in Fig. 16. MONO
Fig. 12 Installation of LVDTs for a MONO, SP1-R, SP3 and b SP1, SP2, and SP3-R specimens.
Fig. 13 Installation of strain gauges on beams and column reinforcing bars (units in mm).
specimen attained the maximum strength at 1.75% drift
ratio. Severe pinching in load–drift relation was observed
due to bond slip of beam longitudinal rebars at 2.75% drift
ratio as seen in Fig. 16a where bond requirements were
satisfied given by the codes (ACI 352 2002; ACI 318 2011).
As the embedment length (equals to column depth) of beam
reinforcing bars within the beam-column joint decreases,
rebar slip will increase and result in a more pinched
hysteretic plot (Moehle 2014). For precast connections, beam
bottom longitudinal rebars were welded and only top rebars
have continuity throughout the joint. Embedment lengths of
top rebars could be increased from column depth (hc) to a
length with the sum of column depth and corbel lengths
(hc ? Lc) which led to an increased energy dissipation
In SP1 specimen, maximum strength was attained at 2.2%
drift ratio in the pull and push directions as shown in
Fig. 16b. 40% degradation in strength for SP1 specimen was
observed due to premature failure of beam welded
longitudinal rebars. SP2 and SP3 specimens showed similar
behavior and maximum strengths were attained at 2.2% drift
ratio as shown in Fig. 16c, d.
In SP1-R specimen, tests were performed up to 5% and
3.5% drift ratios in push and pull directions as shown in
Fig. 16e, respectively. SP1-R specimen showed a more
ductile behavior compare to SP1 specimen. However, shear
failures after 3.5% drift ratio led severe pinching as can be
seen in load–drift relationship. SP3-R specimen attained
maximum strength at 2.2% drift ratio and first degradation in
strength was observed at the second cycle of 3.5 drift ratio as
shown in Fig. 16f.
3.3 Local Response
Beam curvatures, strain in reinforcing bars and shear
deformations were determined for each specimen based on
experimental measurements to obtain local response
quantities. Beam curvatures in connection specimens were
calculated based on the data taken from LVDTs (1 and 2)
mounted on the connection region which includes beam ends
and the joint panel as shown in Fig. 12b. For all specimens,
Fig. 14 Damage patterns of test specimens at each drift ratio and the end of test. a MONO, b SP1, c SP1-R, d SP3, and e SP3-R.
Table 3 Drift ratios corresponding to initiation of each observed damage state.
Flexural crack (%)
Diagonal cracks in Concrete spalling
joint panel (%) (%)
2.2% (2nd cycle)
2.2% (2nd cycle)
3.5% (3rd cycle)
3.5% (2nd cycle)
3.5% (3rd cycle)
Fig. 16 Lateral load–drift ratio relationships for monolithic and precast connection specimens.
moments were determined for Beam 1 element (Fig. 7a) at
the joint face using the data taken from load cell which is
embedded in the pendulum support. Figure 17 shows
moment–curvature relationships for beams of beam-column
Measured strains of beam and column longitudinal rebars
of all specimens are shown in Fig. 18 where strain gauge
locations are given in Fig. 13. Large plastic strains in the
beam longitudinal rebars were developed in the vicinity of
the welded connections as shown in Fig. 18a–e. Drift ratios
at which the strain gauge capacities exceeded the strain limit
of the gauges are also noted in the figures. In SP1-R and
SP3-R specimens, application of unbonded length could be
able to increase the number of strain cycles of welded
longitudinal rebars as well as plastic deformation capacities. On
the other hand, column longitudinal rebars remained nearly
elastic during the tests as shown in Fig. 18f–i.
Shear deformations of beams (Ds) in the vicinity of
connections of SP1, SP2 and SP3-R specimens were calculated
based on the measurements using D1 and D2 transducers
shown in Fig. 12b. Shear deformations are calculated as,
where h is the inclination angle of diagonal LVDT’s (45 )
with respect to horizontal axis with Fig. 19 shows beam
shear force–shear deformation relationships. In SP3-R
specimen which has additional vertical ties and more
frequent transverse reinforcement, shear deformations in beam
elements could be reduced reasonably.
4. Evaluation of Experimental Results
4.1 Lateral Strength and Ductility
Lateral load–drift ratio envelopes of test specimens were
obtained by combining the maximum lateral loads
corresponding to each drift ratio. Component Equivalency
Fig. 17 Moment-curvature relationships of beam-column connection specimens.
Methodology defined in FEMA P-795 (2011) quantifies the
strength, ductility and initial stiffnesses as shown in Fig. 20.
Ultimate drift ratio (HU) corresponds to 80% of the ultimate
strength (QM). Also, effective yield drift ratio (HY,eff) and
initial stiffness (Ki) can be obtained by intersecting the lines
crossing the 0.4QM and QM as shown in Fig. 20.
Figure 21a shows the lateral load–drift ratio envelopes of
test specimens, and normalized strength with respect to yield
strength (Q/Qy) are shown in Fig. 21b. Ultimate strength,
drift ratio and effective ductility of test specimens are
summarized in Table 4. Although the tests were continued up to
higher drift ratios to investigate further failure modes,
ultimate drift ratios (HU [%]) do not correspond to these drifts
where the tests finalized. Ultimate drift ratios corresponding
to 80% of the ultimate strengths (QM) are determined for
both push and pull directions, and the minimum of these
drift ratios are taken into account for ductilities of each test
specimen. In this regard, changing the details for welding of
rebars showed a significant effect on improving the seismic
response. SP1 specimen beams reinforced with relatively
high carbon content bars showed an abrupt decrease in
strength at 2.2% drift ratio. However, SP2 and SP3
specimens showed a gradual strength degradation up to 3.5 drift
Acceptance criteria for moment frames based on
structural testing (ACI 374.1 2005) requires that strength of
specimens at the third cycle of 3.5% drift ratio (Q3.5) should
not be less than 0.75QM. Since in-cycle strength degradation
was caused by abrupt rupture of welded reinforcing bars for
precast specimens, this condition was provided by SP1-R
and SP3-R specimens only as indicated in Table 4—Q3.5
Displacement ductility (le) of a connection specimen is
the ratio of ultimate drift ratio (HU) to the effective yield
drift ratio (He). Displacement ductility of each specimen is
shown in Table 4. SP3-R had higher displacement ductility
among the precast specimens after the achieved
4.2 Lateral Stiffness
Secant stiffness (Ksec) can be calculated as the slope of the
line between the lateral load and corresponding drift ratios as
shown in Fig. 22. ACI 374.1 (2005) requires that secant
stiffness calculated at the third cycle of 3.5% drift ratio
(K3.5) should not be less than 0.05 of the initial stiffness (Ki).
Initial stiffnesses of test specimens are given in Table 5 and
secant stiffnesses at 3.5% drift ratio were provided by all
specimens except for SP1 specimen.
Secant stiffnesses of test specimens were normalized
based on the stiffness at first cycle of 0.15% drift ratio and
shown in Fig. 23. MONO specimen showed faster stiffness
degradation than the precast specimens up to 2% drift ratio.
SP1 and SP1-R specimens showed 58% degradation up to
2.2% drift ratio, however SP1 showed a sudden decrease in
stiffness at 2.2% drift ratio. Besides, SP2, SP3 and SP3-R
specimens showed a gradual decrease in stiffness during the
4.3 Energy Dissipation
Energy dissipation of monolithic and precast specimens
are compared by using relative energy dissipation ratios (bi).
Relative energy dissipation calculated at the third cycle of
3.5% drift ratio should not be less than 12.5% based on ACI
374.1 (2005) criteria. Relative energy dissipation is the ratio
of energy dissipation at each drift to the ideal energy
dissipation represented by the area of the parallelogram shown in
Fig. 22. Relative energy dissipation (bi) can be obtained by
Fig. 18 Measured longitudinal bar strains of beams and columns of connection specimens.
Fig. 19 Beam shear force-shear deformations relationships for SP1, SP2 and SP3-R specimens.
where Ah,i is the area of the closed loop of ith drift ratio, E1i
and E2i are the strengths and h1i and h2i are inelastic drift
ratios in both loading directions.
Figure 24 shows the relative energy dissipation
relationship of test specimens corresponding to third cycle of each
drift ratio. bi values showed a decrease up to 1% drift ratio
inherently because of the minor cracks in the test specimens.
However, yielding of reinforcement and other failure
mechanisms caused an increase in the ratios. All specimens
fulfilled the relative energy dissipation criteria mentioned
4.4 Damage Index
Damage index enables seismic evaluation of existing
reinforced concrete buildings subjected to earthquakes.
Damage indices of test specimens were calculated based on
Fig. 21 a Lateral load–drift ratio envelopes of test specimens and b normalized envelopes of test specimens.
Table 4 Lateral strength and ductility of test specimens.
QM (kN) 0.75 QM (kN) Q3.5 (kN) HY,eff (%)
189.3 142.0 175.5 0.8
183.8 137.9 144.8 0.96
231.6 173.7 140.3 0.86
287.8 215.9 N.A. 1.0
263.2 197.4 210.6 1.1
256.9 192.7 130.6 0.96
249.5 187.1 179.1 0.91
237.5 178.1 161.7 1.16
231.2 173.4 168 0.9
235.8 176.9 154.5 0.9
225.7 169.3 125.7 0.7
228.2 171.1 173.2 1.0
of secant stiffness (Ksec) for test
strength, ductility and energy dissipation characteristics.
Park and Ang (1985) stated damage index (D) in terms of
energy dissipation and displacements as
where QY is the yield strength, EH is the dissipated energy, b
is strength degradation parameter dM and dU are maximum
displacements under earthquake and monotonic loading,
respectively. Ang et al. (1993) proposed the damage index at
collapse state as 0.8.
Craifaleanu and Lungu (2008) derived another damage
index in terms of displacement ductilities as,
D ¼ llU þ b ðlElU 1Þ ð6Þ
Fig. 23 Normalized secant stiffness (Ksec) of test specimens
for corresponding drift ratios.
Fig. 24 Relative energy dissipation ratios (bi) of test
Fig. 25 Damage indices obtained for each test specimen.
Since beam-column connection tests were performed
under cyclic loading, maximum displacements under
monotonic loading (dU) are considered as
where dU,cyc is the ultimate displacement under cyclic
loading conditions based on numerical simulations.
Damage indices of test specimens for the third cycle of
each drift ratio are calculated by Eq. (6) and shown in
Fig. 25. Accordingly, damage index for SP1 specimen has
the highest value due to observed abrupt failure while
unbonded length applied in the SP1-R specimen caused
lower damage index. In addition, with the improvements in
the SP3-R specimen, damage index was determined to have
a lower value accordingly.
In this study, experimental cyclic response of hybrid
(emulative-welded) precast beam- column connections were
investigated. Test specimens were compared in terms of
strength, stiffness and energy dissipation capacities as well
as damage indices.
During the test of monolithic specimen (MONO),
bondslip of beam reinforcing bars after 3.5% drift ratio caused
a pinched behavior in the force–deformation relation and
decreased the energy dissipation. On the other hand,
precast specimens showed increased relative energy
dissipation ratios with the increase of embedment length
of beam top reinforcing bars.
Precast specimens which were tested at the first stage
(without an unbonded length) didn’t show similar
behavior in terms of ductility as the MONO specimen
showed. Strain development of beam welded rebars
played an important role on the overall behavior of
precast connections. SP1 specimen beams reinforced
with relatively high carbon content bars showed an
abrupt decrease in the strength at 2.2% drift ratio.
However, SP2 and SP3 specimens showed a gradual
strength degradation up to 3.5% drift ratio. Obtained
damage indices for SP1, SP2 and SP3 specimens were
close to the index corresponding to collapse state
contrast to MONO specimen.
Unbonding of welded rebars within a sleeve could be
able to decrease strain demands in the vicinity of the
connection. SP1-R specimen showed a higher ductility
than SP1 specimen but it showed severe pinching due to
SP3-R specimen showed an improved seismic behavior
thanks to unbonded length approach and the additional
ties to prevent early buckling of longitudinal bars.
Based on the experimental findings and further
evaluations, revising the relevant requirements in design codes
such as maximum spacing of beam transverse
reinforcement, adopting additional vertical ties and the unbonded
length approach to resist early buckling and rupture of
flexural rebars is proposed.
Experimental study was supported by Turkish Precast
Concrete Association and Dokuz Eylul University Scientific
Research Program under the Grant No. DEU-BAP
2012.KB.FEN.019. The contributions of Dr. Sevket Ozden,
M.Sc. Hakan Atakoy, M.Sc. Gunkut Barka, Dr. Turkay
Baran, Dr. Ozgur Ozcelik and M.Sc. Eng. Umut Yucel to the
experimental studies carried out in the Structural Mechanics
Laboratory at Dokuz Eylul University are gratefully
acknowledged. Tensile tests were performed in Mechanic
Laboratory at Dokuz Eylul University Metallurgical and
Minerals Engineering Department.
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