Seismic Performance of Prefabricated Beam-to- column Joint ...
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Seismic Performance of Prefabricated Beam-to-column Joint with Replaceable Energy-dissipatingSteel HingeLianqiong Zheng ( [email protected] )
Fujian University of Technology https://orcid.org/0000-0002-8996-8151Xiaoyang Chen
Fujian University of TechnologyChanggui Wei
Fujian University of TechnologyGuiyun Yan
Fujian University of Technology
Research Article
Keywords: Prefabricated beam-to-column joint, Replaceable energy-dissipating steel hinge, Prefabricatedsteel tube con�ned joint core, Hysteretic test, Seismic performance
Posted Date: July 24th, 2021
DOI: https://doi.org/10.21203/rs.3.rs-725075/v1
License: This work is licensed under a Creative Commons Attribution 4.0 International License. Read Full License
1
16 June 2021
Dear Editor:
I wish to submit an original paper for publication in Bulletin of Earthquake Engineering, titled
“Seismic performance of prefabricated beam-to-column joint with replaceable
energy-dissipating steel hinge.” The paper was coauthored by Xiao-Yang Chen, Chang-Gui Wei,
and Gui-Yun Yan.
This study presents a novel prefabricated beam-to-column joint for frame and hysteretic tests was
conduct to evaluates the seismic performance and the restorable functional characteristics of the
proposed joints. We believe that our study makes a significant contribution to the literature because
the proposed prefabricated joint providing advantages for precast concrete reinforced frames, such
as complete assembly, damage control, and maintainability of the structure after an earthquake.
This manuscript has not been published or presented elsewhere in part or in entirety and is not
under consideration by another journal. We have read and understood your journal’s policies, and we believe that neither the manuscript nor the study violates any of these. There are no conflicts of
interest to declare.
Thank you for your consideration. I look forward to hearing from you.
Sincerely,
Lian-Qiong Zheng
School of Civil Engineering, Fujian University of Technology
Fuzhou 350118, Fujian Province, People’s Republic of China
Tel.: (+86) 18950383840
Email address: [email protected]
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Seismic performance of prefabricated beam-to-column joint with
replaceable energy-dissipating steel hinge
Lian-Qiong Zheng a,b,*, Xiao-Yang Chen a, Chang-Gui Wei a and Gui-Yun Yan a,b
a School of Civil Engineering, Fujian University of Technology, Fuzhou 350118, P.R. China
(*Corresponding author, E-mail: [email protected])
b Fujian Provincial Key Laboratory of Advanced Technology and Informatization in Civil Engineering, Fuzhou
350118, Fujian Province, P.R. China
Abstract: This study presents a novel energy-dissipating prefabricated joint for connecting beam to
column in a precast frame structure. The joints are characterized by a replaceable steel hinge and a
prefabricated steel tube confined joint core, providing advantages for precast concrete reinforced
frames, such as complete assembly, damage control, and maintainability of the structure after an
earthquake. The hysteretic behavior of the proposed prefabricated joint was studied through two tests.
First, a full-scale prefabricated joint was tested under cyclic loading until failure. On the basis of the
initial test, only four weakened dissipaters of the steel hinges in the prefabricated joint were replaced
and the second test was conducted to investigate the restorable functional characteristics of the
proposed prefabricated joints. For comparison, a reference monolithic joint was also tested. The
experimental results demonstrate that the novel prefabricated beam-to-column joint displayed
excellent hysteretic performance, and corresponding to the monolithic joint, the load-bearing, energy
dissipation, and deformation capacity were improved. The damage of the prefabricated joint was
concentrated on the weakened dissipaters of the steel hinges, indicating that the failure mode and
damage degree of the prefabricated joint can be controlled. In the second test, the prefabricated joint
exhibited similar hysteretic behavior to that of the first test; however, the initial stiffness was slightly
lower. Therefore, the prefabricated joint can meet the replaceability requirement and achieve
satisfactory beam-to-column joint function recovery after an earthquake.
Keywords: Prefabricated beam-to-column joint, Replaceable energy-dissipating steel hinge,
Prefabricated steel tube confined joint core, Hysteretic test, Seismic performance
1
1. Introduction 1
Prefabricated buildings have attracted increased attention in research owing to their advantages of 2
high quality, easy in situ operation, short construction period, and cost savings. Moreover, 3
prefabricated buildings meet the growing demand for building industrialization throughout the 4
world and also satisfy the requirements of environmental friendliness and sustainability. 5
Consequently, prefabricated structures have been extensively applied in building construction [1–5]. 6
Precast concrete frames account for a large proportion of prefabricated buildings owing to their 7
flexibility regarding architectural layout and standardization of the modular components. However, 8
in contrast to conventional cast-in-place structures, a prefabricated building structure demonstrates 9
poor integrity and most prefabricated building structures lose their normal functionality with 10
connection failure due to complex stress conditions, which is important in force transfer between 11
components; in particular, the beam and column. Beam-to-column joints are a vulnerable 12
component and thus it is important to develop good mechanical performance for their use in 13
prefabricated building structures. 14
Therefore, substantial research has been devoted to developing new connections; particularly, 15
beam-to-column joints [6–12] with excellent mechanical behavior for application in prefabricated 16
building structures to achieve good structural integrity. Previous studies indicated that connections 17
between precast frame members could be classified into three main classes: dry connection, wet 18
connection, and hybrid connection. The equivalent monolithic reinforced concrete frame system 19
was developed using cast-in-situ joints or fabricated connections [13–18]. In addition, considerable 20
research has been conducted to better understand the mechanical behavior and evaluate the seismic 21
performance of newly developed connections or precast structures with new connections [7]. 22
2
Restrepo et al. [19] conducted experimental research on four mid-span connections and two 23
beam-to-column connections. It was shown that the connection configurations had a slight effect 24
on the mechanical behavior of the structure. All tested connections exhibited good ductility, 25
energy-dissipation capacity, and bearing resistance and such connections could be constructed to 26
emulate monolithic cast-in-place connections. Zhao et al. [20] investigated the behavior of a precast 27
beam-to-column joint with high-strength concrete, using a full-scale test, and the structure with this 28
new connection demonstrated equivalently desirable seismic performance, such as failure mode, 29
hysteresis behavior, and ductility. Parastesh et al. [21] proposed a novel ductile moment-resisting 30
connection for precast RC frames in seismic regions, which could provide excellent structural 31
integrity and rapid construction. The seismic performance was investigated experimentally and the 32
results demonstrated that the proposed connections could develop adequate flexural strength, as 33
well as higher ductility and energy consumption. Moreover, the failure mode could be improved by 34
concentrating damage to the plastic hinge region. 35
Prestressed steel strands can effectively provide a self-centering capacity for beam-column joints 36
and improve the energy dissipation capacity, without serious structural damage [22–31]. Li et al. 37
[27] researched precast beam-to-column joint subjected to bidirectional lateral loading to assess 38
precast RC frames under the action of an earthquake. The results demonstrated that the prestressed 39
connection exhibited similar performance to the cast-in-place connection. This column remained 40
damage-free during the loading period and an equivalent viscous damping pattern was established, 41
considering the effects of prestress loss and energy dissipater yield. Wang et al. [28] proposed a 42
prefabricated prestressed beam-column joint, which uses replaceable mild steel reinforced bars to 43
provide an energy disappearance capacity and steel strands to provide a self-centering capacity. 44
The effectiveness of the proposed joint was experimentally validated. In addition, the configuration 45
3
was improved based on the results and a parametric study using a calibrated FEA model was 46
conducted to determine the essential parameters the effect of the parameters of interest on joint 47
performance [29]. Through experiments and numerical research, the design specifications to 48
ensure full utilization of the joints were determined. Wang et al. [30,31] developed an all-steel 49
bamboo-shaped energy dissipater and applied it to precast concrete beam-column joints, which 50
played an important role in fusing and protecting the main structure. Through experimental 51
research on five precast concrete connections under cyclic loading, it was found that these 52
connections exhibit good hysteretic performance and self-centering action. 53
Using supplemental energy dissipaters in connection is another research topic aimed at enhancing 54
the energy dissipation capacity and concentrating plastic damage to target members rather than 55
beams or columns, such as top-and-seat angles [32,33], reduced flanges [34], and friction devices 56
[35–37]. To improve the failure mode of the frame structure, a beam-column hybrid joint using 57
flange cover connecting plates for energy-consuming and web connection plates for load 58
transferring [15] is proposed. The efficiency of the hybrid joint is investigated based on 59
experimental research on PC joints and monolithic control connections. Song et al. [23] used bolted 60
web friction devices and a self-centering prestressed connection and adopted them in a 61
moment-resisting frame to reduce residual drift under large drifts and improve the energy 62
dissipation ability with friction damping. Experimental and numerical studies were performed to 63
evaluate its effectiveness. The research results demonstrated that this bolted connection has 64
capabilities of energy dissipation and self-centering comparable to welding connections and 65
simultaneously avoids expensive field welding. Li et al. [38] developed an innovative type of 66
prefabricated beam-to-column joint and investigated the influence of the damper’s geometric 67
dimensions on the hysteretic performance of the prefabricated joint. Similarly, another type of 68
4
damper was proposed by Qi et al. [39] for prefabricated beam-to-column joint, and the seismic 69
performance of the joint were studied as well as the design procedure was proposed. 70
Although previous studies have demonstrated that the behavior of prestressed joints or hybrid 71
connections is comparable to that of monolithic connections [40,41], there are still issues, such as 72
the difficulty in constructing connections between framing components, the feasibility of damage 73
or failure mode control, and the potential for replacing the damaged energy dissipaters. Therefore, 74
according to the present research, a new prefabricated beam-to-column joint is developed to 75
mitigate the unsolved issue. Pseudo-static research was conducted on a novel prefabricated joint 76
and the failure mode, load-bearing capacity, stiffness, ductility, energy dissipation, and 77
earthquake-resilience were analyzed. In addition, the dissipaters of the steel hinges in the 78
prefabricated joint were replaced and a second hysteretic test was conducted to study the restorable 79
characteristics of the proposed prefabricated joints. 80
81
2. Mechanism of novel prefabricated joint 82
Based on the combination of load-carrying and energy-dissipation elements, an innovative type of 83
prefabricated beam-to-column joint was developed, according to the schematic illustration in Fig. 1. 84
The new beam-to-column joint is characterized by replaceable steel hinge and prefabricated steel 85
tube confined joint core, providing advantages, such as complete assembly, damage control, and 86
maintainability of the structure after an earthquake. 87
88
2.1 Replaceable energy-dissipating steel hinge
As described in Fig. 1a and b, replaceable steel hinges are set at the ends of the precast beam 89
adjacent to the column. I-shaped steel with an endplate is embedded at the precast beam end. In 90
addition, I-shaped steel is welded at both sides of the joint core steel tube. Therefore, the precast 91
5
beam can be bolted to the column via a steel hinge. The steel hinge is composed of two types of 92
connected components, replaceable upper and lower energy dissipaters and the pin shaft connection, 93
as shown in Fig. 1c. Low-yield point steel (LYP) is cut into a dog-bone shape to reduce geometry as 94
described in steel design codes, such as FEMA 350–351 [42,43] and EC8 [44], and an arc-shaped 95
stiffener is welded under the LYP plate to avoid premature buckling. The LYP plate, arc-shaped 96
stiffener, and end plates formed an energy dissipater, which was used to transfer the bending 97
moment between the beam and column, and also acted as a fuse to dissipate energy and concentrate 98
plasticity under earthquake excitations. Weakened areas are easily formed due to the low yield 99
strength and reduced section of the LYP plate; therefore, the damage position and damage degree 100
can be controlled. Owing to the excellent seismic behavior of LYP steel [45], the LYP plate will 101
concentrate most of the plasticity and damage, maintain the main structural members, such as 102
beams and columns, and remain elastic. The pin shaft connection of the steel hinge contains left and 103
right lugs, a high-strength pin, and end-plates. The pin hinges the lugs and supplies the shear 104
capacity in the steel hinge. The steel hinge adequately rotates around the pin, which can outward the 105
plastic hinge away from the beam-column interface and shift into the weakened region of the steel 106
hinge. This can improve the uncertainty of the location of plastic hinges in traditional reinforcement 107
concrete frames owing to their limited rotation ability. 108
Steel hinges are used to connect the precast beam and column with high-strength bolts, which are 109
easy to manufacture, disassemble, and replace. Moreover, endplates of the energy dissipaters and 110
endplates of lugs are independent, as shown in Fig. 1c, allowing for the replacement of the 111
dissipaters instead of the entire steel hinge after an earthquake. 112
6
113
(a) Precast concrete frame (b) External joint (c) Steel hinge 114
Fig. 1. Description of the prefabricated joint for a precast concrete frame with replaceable 115
energy-dissipating steel hinge 116
117
2.2 Prefabricated steel tube confined joint core
For the newly developed prefabricated beam-to-column joint, a steel tube was employed to confine 118
the core area to realize the seismic design concept commonly referred to as a strong connection, as 119
well as to connect the upper and lower precast column segments. As shown in Fig. 2, the proposed 120
prefabricated steel tube confined joint core is comprises a steel tube, joint core concrete, internal 121
diaphragms, connected components, and ducts for the insertion of reinforcements and grout pouring. 122
The joint core concrete is confined by the steel tube, improving its shear resistance. Thus, shear 123
failure in the joint core area can be avoided. Two internal diaphragms are welded in the steel tube to 124
transfer the horizontal force in the core area, which is equivalent to the continuous longitudinal 125
reinforced steel bar of the beam passing through the core area utilized in the cast-in-place 126
beam-to-column joint. 127
Connected component: Pin Shaft Connection
Lugs
Pin
Endplate Beam
Precast RC column Precast RC beam
Connected component:
Replaceable energy dissipater
(Fuse, yield firstly, energy
dissipation, damage concentration)
Beam end plastic
steel hinge
Co
lum
n
Prefabricated steel tube confined joint core
Steel hinge
Stiffener
Low yield point
steel (LYP) plate
Endplate
Embedded I-shaped steel in the beam
I-shaped steel welded at
joint core steel tube
7
128
Fig.2. Proposed prefabricated steel tube confined joint core 129
The connection system between precast columns in this research is based on the use of precast 130
columns, with sleeves encased on the lower end and longitudinal reinforcements protruding from 131
the upper end. For the column-to-column connection assembly, as displayed in Fig. 3, the lower 132
precast column is placed in position. Then, the prefabricated joint core is lifted, aligned with the 133
lower column, and lowered to insert the protruding reinforcements of 134
the lower column through the corrugated steel ducts of the joint core. High-strength grout is then 135
poured to fill the ducts and create a layer at the interface between the lower column and the joint 136
core. Finally, the upper precast column is lifted and positioned to insert the protruding 137
reinforcements of the lower precast column into the sleeves of the upper precast column. The 138
sleeves were grouted to achieve continuous longitudinal reinforcement of the column and the 139
interface between the upper precast column and prefabricated joint core was simultaneously 140
grouted. Both ends of the precast column are concrete rough surfaces to ensure the transfer of shear 141
force at the interface. 142
143
component: Connect with
steel hinge at beam end
Steel tube
Joint core concrete
Ducts for insertion
of protruding bars
Upper internal diaphragm
Duct for grout
pouring
Lower internal diaphragm
8
144
Fig.3. Connection system between precast columns 145
3. Experimental Program 146
3.1 Test specimens
Two quasi-static tests were conducted to investigate the hysteretic behavior of the novel 147
prefabricated beam-to-column joint. First, a prefabricated joint was tested under cyclic loading until 148
failure. According to this previous test, only four weakened steel hinge dissipaters were replaced 149
and the second test was conducted to investigate the restorable functional characteristics of the 150
proposed prefabricated joints. The specimens of the prefabricated joint under the two loadings were 151
labeled as PJ-1 and PJ-2, respectively. Finally, a cast-in-place monolithic joint specimen (labeled 152
MJ) was also tested to estimate the seismic behavior of the proposed prefabricated joint. 153
This specimen adopts an interior beam-to-column joint idealized from the prototype structure; the 154
beam has a span of 4000 mm, and the length of the column is 3040 mm. This prefabricated joint was 155
composed of precast RC columns, precast RC beams, a steel tube confined joint core, and an 156
energy-dissipating steel hinge with weakened flanges. The detailed configuration of the 157
Protruding
reinforcement
layer filled at the
interface
Lower precast column
Grouted corrugated
steel ducts
Grout inlet
layer filled at the
interface Lower precast column
Grouted corrugated
steel ducts
Upper precast
column
layer filled at
the interface
Grout outlet
Sleeves
Connect with steel
hinge at beam end Pouring grout
9
prefabricated joint and geometrical dimensions of the steel hinge and steel tube confined joint core 158
are presented in Fig. 4. The section of the precast beams was 250 mm × 550 mm (width × height) 159
and the section width of the square precast column was 400 mm. The precast column and beam 160
were reinforced with 12 longitudinal bars 22 mm in diameter and eight bars that were 18 mm in 161
diameter, respectively. In addition, 8 mm diameter hoops were provided and the stirrup spacing was 162
150 mm. For comparison, as shown in Fig. 5, the geometry and reinforcement of the monolithic 163
joint were identical to those of the prefabricated joint. 164
165
166
Section (1-1) 167
Fig. 4. Prefabricated specimen configuration (unit: mm) 168
250
4000
200 125 125 Replaceable energy dissipater (Weakened)
950 150 50
450 50
400 150 50
450 50
150 950 150
75
400
400
1370
1170
500
1800 1800
Welded to flange of steel beam Internal Connected
400
Grouted sleeve
Section (3-3)
Endplates and Bolts
400
400
12 22
8@150
Endplate
2
2
3 3
1 1
①Embedded I-shaped steel
550
250 60 60 60 35 35
M20
M30
25
25
85
105
120
105
85
250
550
8 18
Section (2-2)
8@150
Steel hinge
17
5
175
200
550
25
25
85
65
120
65
85
40
40 R9050 R20
225400
50
180
225
R60
60
10
105
②
Steel tube
100 100 100 50 400
100
50
100 5
0
Internal diaphragm
R50
R20
10
0
10
10
250
10
Section of ① and ②
430
550
10
169
Fig. 5. Monolithic specimen configuration (unit: mm) 170
3.2 Specimen construction
The fabrication procedure of this prefabricated joint began with the production of the upper column, 171
lower column, beams, steel tube confined joint core, and energy-dissipating steel hinges, as shown 172
in Fig. 6. The formwork for the precast columns and precast beams was prepared using plywood. 173
Then, the reinforcements and grouted sleeves of the columns were installed, and the upper and 174
lower flanges of an I-shaped steel beam with a length of 300 mm were welded with the 175
corresponding longitudinal reinforcements of beams. Subsequently, concrete with a strength grade 176
of C55 was poured for columns and beams and cured for 14 days. A steel tube was welded using 177
four plates measuring 400 × 500 mm2 for confining the core area. Two internal diaphragms were 178
welded to the steel tube at the position aligned with longitudinal reinforcements of the precast beam 179
400
Section (3-3) Section (4-4) Section (1-1)
250
550
8 18
8@100
250
550
8 18
8@150
12 22
8@100
400
400
Section (2-2)
400
1800 1250 550
1800 1250 550
3
3
550
945
12 22
8@150
400
400
400
745
1 1
2 2
4
4
11
to ensure a reliable bending moment at the beam ends. After two I-shaped connecting beams were 180
welded to the steel tube, concrete was cast into the joint core between two internal diaphragms. 181
182
Fig. 6. View of the prefabricated specimen series for the (a) reinforcement and grouted sleeve of the 183
upper precast column, (b) lower precast column, (c) precast beams, (d) prefabricated steel tube 184
confined joint core, and (e) energy-dissipating steel hinge 185
186
When the concrete strength of the columns, beams, and joint core reached the hoisting allowable 187
strength, the top surfaces of the lower column and bottom surface of the upper column were 188
roughened, as shown in Fig. 6b, to transfer the shear force between the interfaces of the upper 189
column, joint core, and lower column. Then, the prefabricated components were assembled 190
according to the following steps: (1) presented in Fig. 7a, hoist the prefabricated steel tube confined 191
joint core to make the protruding longitudinal reinforcements in the lower precast column through 192
the reserved hole formed by the steel syphon bellows in the joint core. Then, the prefabricated joint 193
core was placed on the top of the lower column. (2) High-performance cement paste was grouted to 194
the duct between the bellows and reinforcements, as shown in Fig. 7b, to ensure the stability of the 195
reinforcements and connect the top surface and joint core. (3) The upper precast column was 196
installed at the top of the joint core (Fig. 7c). (4) The sleeves of the upper precast column were 197
grouted to connect the longitudinal reinforcements of the lower and upper columns and the bottom 198
Grouted sleeve a b Steel tube d
c Roughen surface
I-shape steel beam (150mm embedded
and 150mm for connecting)
e
Internal diaphragm
I-shaped connecting beam
Lugs
Pin
Dissipater Pouring hole
Rebar through hole
12
surface of the upper column and joint core. As shown in Fig. 7d, a high-performance cement paste 199
was grouted from the bottom and flowed out of the top to ensure that the cavity of the sleeve was 200
filled with paste. (5) The steel hinge was connected to the precast beam using high-strength bolts 201
(Fig. 7e). (6) Two precast beams and steel hinges were connected to each side of the joint core (Fig. 7f). 202
203
Fig. 7. Assembly process of the prefabricated joint: (a) assemble the prefabricated steel tube 204
confined joint core and lower precast column, (b) duct grouting of the joint core, (c) install the 205
upper precast column, (d) grout the sleeves of the upper precast column, (e) install the 206
energy-dissipating steel hinge, and (f) connect the precast beams and columns. 207
208
3.3 Material properties
LYP was applied to a replaceable energy dissipater. The other steel components, including the lugs, 209
I-shape steel, steel tube, and internal diaphragms of the beam and column connector, were made of 210
Q355. HRB400 was used as reinforcements in the precast beams and columns. The material 211
properties of the steel plates and bars were obtained by tensile tests that were based on 212
GB/T228.1-2010 [46]. The determined yield strength (fy), yield strain (εy), ultimate strength (fu), 213
elastic modulus (Es), and Poisson’s ratio (μs) values are listed in Table 1. High-strength bolts (Class 214
10.9) were employed for connections between the prefabricated joint core, steel energy-dissipating 215
hinge, and precast beam. 216
a b c
f d e
13
All specimens were cast using concrete with identical mix proportions. The cubic compressive 217
strength (fcu) and elastic modulus (Ec) of the concrete were measured by a cube with dimensions 218
of 150 mm × 150 mm × 150 mm and a prism with dimensions of 150 mm × 150 mm × 300 mm, 219
respectively. At the testing days, fcu and Ec of the concrete were 56 MPa and 35400 MPa, 220
respectively. 221
222
Table 1 Material Properties of the steel plates and bars 223
Type
Thickness
(Rebar diameter)
(mm)
fy (MPa) fu (MPa) y () Es (MPa) s
Low-yield-point steel 10 269.8 373.6 1659 200334 0.284
Steel plate (Q355) 10 374.2 489.7 2322 193229 0.273
Longitudinal reinforcement of the columns 22 414.9 563.7 2527 196966 0.294
Longitudinal reinforcement of the beams 18 419.9 558.5 2446 199576 0.283
Stirrups of the columns and beams 8 432.2 572.5 2612 200566 0.287
224
3.4 Test setup and loading scheme
Hinge support boundary conditions are set for the column bottom and beam ends to simulate 225
infection points in a frame. To reflect the actual working conditions and account for the second-order 226
influence in the frame, a constant axial load was applied to the free top of the column during the 227
lateral cyclic loading for all specimens. First, an axial compression of 1600 kN (an axial load ratio of 228
approximately 0.3) was applied and maintained constant using a jack with a capability of 2000 kN. 229
This jack was connected to a rigid reaction beam by a rolling support to ensure that the compression 230
on the column was concentric. The cyclic load was provided by a hydraulic actuator with an ability of 231
±500 kN. The test setup is shown in Fig. 8. 232
The tests were conducted under force and displacement control following ACT-24 [47]. In the force 233
control phase, load levels of 0.25 Puc, 0.5 Puc, and 0.7 Puc were selected, where Puc denotes the 234
predicted lateral ultimate strength by finite element analysis. When the yield strain was observed, the 235
14
displacement control was then obtained at increments of Δy, 1.5Δy, 2Δy, 3Δy …, where Δy represents 236
the yielding displacement. The cyclic loading scheme of the lateral load is shown in Fig. 9. Two 237
cycles and three cycles at each level were conducted for load and displacement control phase, 238
respectively. 239
240
Fig. 8. Photo of test set-up 241
242
243
Fig. 9. Cyclic loading scheme of the lateral load 244
245
During the test, the lateral load-displacement (P-Δ) hysteretic curves were automatically recorded 246
by the loading system. Four extensometers (Nos. 1 – 4 in Fig. 10a) were set at the upper and lower 247
flanges of the steel hinge to indirectly measure the rotation of the steel hinges. The shear drift of the 248
joint core was measured using two extensometers (No. 5 and 6 in Fig. 10a) installed along the 249
-6-5-4-3-2-10123456
0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24
Load
(P
/Puc)
Cycle number
Rolling support
Hydraulic actuator Hydraulic jack
Hinge Support
Upper precast column
Energy-dissipating steel hinge
Precast beam
Lower precast column
Steel tube confined joint core
Dis
pla
cem
ent(
Δ/Δ y)
Force control Displacement control
0.25Puc 0.5Puc 0.7Puc 1.0y 1.5y
2.0y 3.0y
4.0y
15
diagonal lines. Displacement transducers 1,2, and 3,4 were set at the bottom of the upper precast 250
column and the top of the lower precast column, respectively, to evaluate the rotations of the 251
columns. The ranges were ± 200 mm for the displacement transducers and ± 50 mm for the 252
extensometers. 253
For the prefabricated specimens, as presented in Fig. 10b, a total of 13 strain gauges and two strain 254
rosettes were used to obtain the strains in energy dissipaters and the lugs of the steel hinge, 255
respectively. Strain gauges mounted on the longitudinal reinforcements and the steel tube of the 256
joint core are shown in Fig. 10c. For the monolithic specimen, strain gauges were set at a distance of 257
100 mm from the joint core for the longitudinal reinforcement in the columns and beams. 258
259
(a) Transducer location 260
261
(b) Strain gauge layout of the energy-dissipating steel hinge 262
263
Strain gauge
1# 2# 3#
4# 5#
175 175 50 50
450
500
25
25
8# 9# 10#
11# 12# 13#
250
450 125 125 200
50
50
50
50
50
50 50 50 50 50 50
1#
75
3# 5# 6# 7# 4#
75
2#
B
A 45
45
45
45
Extensometers 5, 6 Extensometer 3 Extensometer 1
Extensometer 2 Extensometer 4
Displacement transducer 4 Displacement transducer 3
Displacement transducer 1 Displacement transducer 2
16
(c) Strain gauge layout of the longitudinal reinforcement and confined steel tube of the prefabricated 264
specimens 265
Fig.10. Arrangement of the instrumentation 266
267
4. Test results and analysis 268
4.1 Test observations
4.1.1 Initial loading test for the prefabricated specimen (PJ-1) 269
During the first load step of the prefabricated specimen, specimen PJ-1 experienced an approximate 270
elastic deformation at the force control stage. Yielding occurred on the flange near the maximum 271
weakened position of the replaceable energy dissipaters in steel hinge, at a load of 100 kN 272
(approximately 0.5Puc), and the corresponding displacement of column top was 9.6 mm. After that, 273
displacement control loading was applied based on Δy = 10 mm. 274
When the displacement reached 1.5Δy, four longitudinal shear sliding cracks appeared in the precast 275
beam along the interface between the flanges of the embedded steel and the nearby concrete. The 276
length of the cracks was similar to that of the embedded part of the I-shaped steel. A flexural crack 277
through the bottom of the left beam was observed at the same time, which was approximately 150 278
mm away from the concrete edge of the precast beam. Under reversed loading and the second cycle 279
of 1.5Δy, symmetric flexural cracks at the top and bottom of the beams adjacent to the column were 280
observed. 281
New flexural cracks were observed approximately 150 mm from the initial flexural crack as the 282
displacement increased to 2Δy and further outward flexural cracks developed at 3Δy. In addition, the 283
cracks developed from the bottom and top to the sides of the beam and inclined cracks subsequently 284
formed. Slight cracks appeared in the lower precast column near the joint core at a displacement of 285
4Δy; however, no cracks appeared in the upper precast column owing to the reinforcement of the 286
grouted sleeve during the entire loading process. 287
17
288
Slight local buckling on the energy dissipation hinge in the compression flange was observed at an 289
incremental displacement of 5Δy. Subsequently, the precast beam and column developed no further 290
cracks and the plastic deformations were mainly concentrated in the energy dissipaters of the steel 291
hinge. As the displacement increased, local buckling of the steel hinge became apparent. Specimen 292
PJ-1 attained forward and reversed peak loads at 7Δy and -8Δy, with values of 230.08 kN and -229.02 293
kN, respectively. Then, the load-bearing capacity of the specimen deteriorated owing to further 294
loading. A slight crack was observed in the upper tension flange of the right beam in the first cycle of 295
10Δy, and the fracture developed at the second cycle of 10Δy. Then, the specimen failed. The final 296
failure appearance of specimen PJ-1 is shown in Fig. 11. During the loading process, the strains for 297
the longitudinal reinforcements in the precast beam and column were in the elastic range. 298
299
(a) Finial failure appearance 300
Fracture
Flexural
crack
Local buckling
Local buckling Local buckling
18
301
(b) Crack development of the left precast beam (c) Crack development of the right precast beam 302
Fig.11. Crack development and failure mode of specimen PJ-1 303
304
4.1.2 Second loading test for the prefabricated specimen (PJ-2) 305
Based on the previous test, four damaged dissipaters of the steel hinges were replaced, forming 306
specimen PJ-2 and the second loading test for the prefabricated specimen was conducted under the 307
same scheme as the first loading test. The replacement process is shown in Fig. 12. During the 308
initial stage of loading, there was no new crack development and specimen PJ-2 was in the elastic 309
stage. New cracks at the end of the shear cracks on the beam sides were first observed at 2Δy, 310
following which new crack propagation was rare. The crack development in the precast beams of 311
specimen PJ-2 is marked by the red line, as shown in Fig. 13a. 312
Similar to the first test, slight local buckling on the compression dissipaters of the steel hinge was 313
observed at 5Δy. Subsequently, the precast beam and column developed no further cracks when the 314
plastic deformations were centralized in the dissipaters. In addition, the local buckling of the steel 315
hinge became more obvious upon further loading. Specimen PJ-2 reached its peak load at 7Δy. The 316
values were 225.70 kN and -233.41 kN for the push and pull directions, respectively, which is 317
similar to the load-bearing capacity of specimen PJ-1. The failure area of specimen PJ-2 was in the 318
energy dissipaters of the steel hinge as well as that of specimen PJ-1. At the third cycle of 10Δy, the 319
lower tension dissipater of the right steel hinge fractured and the test was terminated. The 320
deformation development of the steel hinges is shown in Fig. 13b and c. 321
Shear sliding crack
Initial flexural crack Flexural cracks
Shear sliding
cracks
Inclined cracks
19
322
Fig.12. Process of replacing the left energy-dissipating steel hinge 323
324
(a) Crack development of the precast beams 325
326
(b) Failure development characteristic of the left energy-dissipating steel hinge 327
328
(c) Failure development characteristic of the right energy-dissipating steel hinge 329
Fig.13. Crack development and failure mode of specimen PJ-2 330
331
During the two tests for the prefabricated specimens, no other visible buckling or changing 332
phenomenon was observed. The confined steel tube was removed after testing specimen PJ-2 for 333
direct observation of the core concrete. As shown in Fig. 14, only a slight crack was found in the 334
core area, indicating that the novel prefabricated joint met the seismic design concept commonly 335
referred to as “strong connection”. 336
New cracks New cracks
5y 7y 10y
Fracture Severe buckling
Slight buckling
5y 7y 10y
Severe buckling Slight buckling
20
337
Fig.14. Concrete in the core area of the joint 338
4.1.3 Monolithic specimen (MJ) 339
In the monolithic specimen, the first cracks occurred near the beam and column interface at a load 340
of 80 kN (0.5Puc) and some flexural cracks were located at the bottom and top of the beams. The 341
corresponding displacement was 7.6 mm. Then, displacement control loading was applied based on 342
Δy = 7 mm. 343
During the displacement control phase, the flexural cracks in the beams developed from the 344
position close to the core area to the end; then, the inclined crack developed into a through-crack at 345
a displacement of 1.5Δy. When the displacement attained 2Δy, a horizontal flexural crack was found 346
in the column at a position close to the joint core. Upon further loading, no cracks developed in the 347
beam. The occurrence and development of cracks were concentrated in the joint core area. 348
Furthermore, oblique shear cracks were formed at an angle of approximately 38 – 45 ° in the 349
horizontal direction in the joint core area at 3Δy and specimen MJ attained maximum strength. 350
Some concrete gradually began to crush and fall off within the four joint core area corners at 5Δy 351
and this became more serious and extended to the upper column at incremental displacements of 352
5Δy – 7Δy, which caused exposure of the reinforcements. After the peak load, as shown in Fig. 15, at 353
10Δy, the lateral load reduced to 85 % of the peak load, the monolithic joint failed by shearing at the 354
joint core, and the test was terminated. 355
A slight crack
21
356
Fig.15. Failure mode of specimen MJ 357
During testing, the relationship between the precast columns and precast beams was reliable and the 358
prefabricated specimens and monolithic specimen exhibited good overall mechanical performance. 359
Compared with the monolithic joint, the plastic hinge was forced outward from the column face, 360
and damage was controllable in the prefabricated joint owing to the steel hinge. The replaceable 361
energy dissipater made of low-yield-point steel entered the plastic stage first and dissipated the 362
energy. Therefore, the crack development and damage of the precast RC beams, precast RC 363
columns, and joint core area were efficiently controlled. There were no significant differences in the 364
failure mode and load-bearing capacity of the prefabricated specimens under the first and second 365
loading times (specimens PJ-1 and PJ-2). Thus, the hysteretic performance of the proposed novel 366
prefabricated joint can be restored by replacing only the energy dissipater. 367
4.2 Lateral load-displacement (P-Δ) hysteretic curves
The P-Δ hysteretic curves of all specimens are shown in Fig. 16. The hysteretic curves of all 368
specimens approached straight lines and the specimens experienced elastic deformation during the 369
initial loading stage. For prefabricated specimens PJ-1 and PJ-2, the slope of the hysteretic curves 370
decreased slightly as energy dissipater in the steel hinge yielded while the strength increased 371
22
consistently owing to the stress strengthening of the dissipater, and residual displacement was 372
observed when unloading. The P-Δ hysteretic curves of the prefabricated joints are plump in shape 373
and have no obvious pinching effect, as shown in Fig. 16a, indicating excellent energy dissipation 374
capacity. When the lateral displacement reached 80 mm — i.e., that the drift ratio reached 2.5% — 375
the load capacity decreased owing to the local buckling of the dissipaters. When the prefabricated 376
specimens failed, the drift ratio reached 3.33 %. Moreover, the P-Δ hysteretic curves of PJ-1 and 377
PJ-2 coincided, indicating that the seismic behavior of the novel prefabricated joint can be restored 378
by replacing only the energy dissipater. 379
380
The P-Δ hysteretic curves of the monolithic joint exhibited significant pinching in the middle, as 381
shown in Fig. 16b, because specimen MJ failed in a shear-dominant mode in the joint core area. 382
After the peak load, the strength decreased gradually owing to the concrete crushing and spalling. A 383
comparison of all the P-Δ hysteretic curves revealed that the seismic performance, such as the 384
load-bearing capacity, energy consumption, and ductility, of the designed novel prefabricated joint 385
was better than that of the monolithic joint. 386
387
(a) Prefabricated joints (b) Monolithic joint 388
Fig.16. P- hysteretic curves of all specimens 389
4.3 P-Δ envelope curves
-300
-200
-100
0
100
200
300
-150 -100 -50 0 50 100 150
P(k
N)
Δ (mm)
PJ-1
PJ-2
-300
-200
-100
0
100
200
300
-150 -100 -50 0 50 100 150
P(k
N)
Δ (mm)
MJ
23
Fig. 17 shows the P–Δ envelope curves of all specimens. The yield load (Py) and yield displacement 390
(Δy) were determined by the geometrography method, as presented in Fig. 18 in [48]. The load and 391
displacement at the peak point of the P-Δ envelope curve were labeled as Pmax and Δmax, 392
respectively, and the failure load (Pu) and corresponding displacement (Δu) were determined when 393
the load decreased to 85 % of Pmax. The displacement and load at the points of yield, peak, and 394
failure are listed in Table 2 for all specimens. 395
In Fig. 17, the initial stiffness of specimen PJ-2 was slightly lower than that of PJ-1; this is because 396
of the existence of cracks in the PJ-2 specimen precast beams and columns from the beginning, as 397
only the dissipaters were replaced after the first load. Thus, Δy of specimen PJ-2 was slightly higher 398
than that of specimen PJ-1. After the yield point, the difference between the P and Δ envelope 399
curves of specimens PJ-1 and PJ-2 can be considered negligible. The load-bearing, energy 400
dissipation, and deformation capacities can be fully restored by replacing the dissipaters of the steel 401
hinges. 402
As can be seen from Table 2, the load-bearing capacities of prefabricated specimens PJ-1 and PJ-2 403
were 48 % and 49 % higher than that of the monolithic specimen MJ, respectively. This is because 404
specimen MJ failed in a shear-dominant pattern in the joint core area while the core area was 405
enhanced by a steel tube for the prefabricated specimens. Moreover, plastic hinges were bound to 406
occur at the beam ends, which protected the precast beam and column from damage. 407
24
408
Fig.17. Skeleton curves of all specimens 409
410
Fig.18. Determination of the yield point, peak point, and ultimate state of the specimens 411
412
Table 2 Characteristic values of P-Δ envelope curves and ductility coefficients for all specimens 413
Specimen Yield point Peak point Failure point Ductility
coefficient μ
Average
of μ Δy (mm) Py (kN) Δmax (mm) Pmax (kN) Δu (mm) Pu (kN)
PJ-1 17.21 161.54 67.18 230.08 92.21 195.50 5.35
5.79 -16.21 -154.65 -78.48 -229.02 -100.81 -197.56 6.22
PJ-2 26.18 152.91 78.56 225.70 100.83 205.10 3.85
3.99 -24.39 -166.85 -66.88 -233.40 -100.81 -201.80 4.13
MJ 13.42 99.96 22.41 149.20 72.32 126.82 5.38
5.67 -14.27 -102.23 -22.41 -162.40 -85.11 -138.04 5.96
414
4.4 Stiffness degradation
-300
-200
-100
0
100
200
300
-120 -80 -40 0 40 80 120P
(kN
)
(mm)
PJ-1
PJ-2
MJ
Py
Pmax
Pu (0.85Pmax)
y
u
max
P
Yield point
Peak point Failure point
O
25
The average secant stiffness is used as the stiffness of the specimens under different loading levels, 415
and the relative stiffness of the i-th average secant stiffness Ki is defined as [49]: 416 𝐾𝑖 = |+𝑃𝑖|+|−𝑃𝑖||+𝛥𝑖|+|−𝛥𝑖| , (1) 417
where Pi and i are the peak load and lateral displacement, respectively, under the i-th /y and ‘+’ 418
and ‘-’ represent the positive and negative directions, respectively. 419
Fig. 19 compares the Ki – Δ/Δy curves of all specimens. For all specimens, the stiffness decreases as 420
the lateral displacement increases. The stiffness under the first loading cycle for the cast-in-place 421
monolithic joint is approximately 23.1% higher than that of prefabricated specimen PJ-1, 422
demonstrating that the steel hinge connection results in a slight deterioration of the integrity of the 423
beam. However, specimen MJ showed a more severe rate of stiffness degradation. Therefore, the 424
stiffnesses of specimens MJ and PJ-1 were similar at 1.5 Δy. Subsequently, the stiffness of specimen 425
PJ-1 was higher. This is because the cracks occurred and extended continuously during the entire 426
loading process for specimen MJ. In the prefabricated joint, after a certain displacement increment, 427
the damage was concentrated in the steel hinge dispersers, which led to no further cracks in the 428
precast column and beam. 429
From Fig. 19, the initial stiffness of the restored prefabricated specimen PJ-2 is 30 % lower than that 430
of PJ-1, as mentioned in Section 4.3; this is due to the existence of cracks in the precast beams 431
before loading. The two prefabricated specimens, PJ-1 and PJ-2, experienced similar stiffness 432
degradations because the degradation was approximately due to the yield of the dissipaters. 433
26
434
Fig.19. Stiffness reduction curves of the specimens 435
4.5 Strength degradation
Fig. 20 shows the strength degradation of λ2 and λ3 of the specimens as a function of the lateral 436
displacement, where λ2 and λ3 are the strength degradation coefficients of the second and third 437
cycles, respectively, at the same loading level. For prefabricated specimens PJ-1 and PJ-2, λ2 and λ3 438
were stable at approximately 1.0, with a small jitter before fracturing of the dissipater, indicating 439
that the novel prefabricated joint had an excellent load-bearing capacity under cyclic loading. The 440
strength degradation of the monolithic joint occurred earlier than that of the prefabricated joint and 441
the rate of strength degradation was faster; λ2 and λ3 were approximately 0.85 and 0.90, respectively, 442
for specimen MJ. In summary, the strength degradation of MJ is significant. 443
444
(a) 2 (b) 3 445
Fig.20. Strength degradation of the specimens 446
447
4.6 Ductility and energy dissipation
0
2
4
6
8
10
12
0 2 4 6 8K
i(1
03kN
/m)
/y
PJ-1
PJ-2
MJ
0.4
0.6
0.8
1
1.2
-120 -90 -60 -30 0 30 60 90 120
2
(mm)
PJ-1
PJ-2
MJ
0.4
0.6
0.8
1
1.2
-120 -90 -60 -30 0 30 60 90 120
3
(mm)
PJ-1
PJ-2
MJ
27
Following the definition by Han et al. [48], the displacement ductility coefficient (μ) of all 448
specimens is determined by μ = Δu/Δy and listed in Table 2. From Table 2, the average of the active 449
and passive failure displacements for specimens PJ-1 and PJ-2 are 96.51 mm and 100.82 mm, 450
respectively, and the corresponding drift ratios are 3.22 % and 3.36 %, respectively, indicating the 451
excellent deformation ability of the prefabricated joint. The displacement ductility coefficient of 452
PJ-2 decreased by approximately 31 % compared to that of specimen PJ-1, as previously mentioned. 453
This is because the Δy of PJ-2 was slightly higher than that of PJ-1. For specimen MJ, which is 454
cast-in-place, Δy and Δu are 13.85 mm and 78.72 mm, respectively; i.e., significantly lower than 455
those of specimen PJ-1 and PJ-2. In addition, the μ value of specimen PJ-1 was higher than that of 456
specimen MJ. Thus, the deformation capacity of the prefabricated joint was improved. 457
458
The cumulative hysteretic energy (Ep), calculated based on the area enclosed by the hysteretic 459
hoops from the P﹣ hysteretic curves, and the equivalent hysteretic damping coefficient (ζeq), 460
determined according to Fig. 21 shown in [49], was employed to estimate the energy consumption 461
capacity of the joints. The equivalent hysteretic damping coefficient can be expressed as: 462
ζeq= 12π· S(ABC+CDA)
S(OBE+ODF), (2) 463
where SABC+CDA is the hysteresis loop area, and SOBE+ODF is the area of triangle OBE and ODF. 464
28
465
Fig.21. Definition of the equivalent hysteretic damping coefficient 466
Fig. 22 and 23 show the Ep﹣curves and ζeq﹣curves for each specimen. The calculated 467
values of the cumulative hysteretic energy and equivalent hysteretic damping coefficient once the 468
lateral load reduced to 85 % of the ultimate strength — i.e., Ep,u and ζeq,u, respectively — are 469
listed in Table 3. Evidently, Ep increases with increasing displacement. In contrast to monolithic 470
joint MJ, energy dissipation capacity of prefabricated joints PJ-1 and PJ-2 clearly increased owing 471
to the superior plastic energy dissipation capacity of the steel hinge. Here, Ep,u of joints PJ-1 and 472
PJ-2 were higher than that of joint MJ by a factor of approximately 2 and 1.8, respectively. The 473
equivalent hysteretic damping coefficient of the joint MJ developed rapidly during the initial 474
loading stage owing to the serious cracking of the joint core area. However, after the displacement 475
reached 30 mm, the ζeq of the prefabricated joints were higher than that of monolithic joint. The 476
values of ζeq,u for joint PJ-1 and PJ-2 increased by 138 % and 100 %, respectively, compared to 477
joint MJ. From the above, it can be seen that the prefabricated joint provides excellent energy 478
consumption ability through the use of steel hinges. In addition, as shown in Fig. 22 and Fig. 23, 479
the agreement between the cumulative hysteretic energy curves of joint PJ-1 and PJ-2, as well as 480
the similar equivalent hysteretic damping coefficient curves between joint PJ-1 and PJ-2, reveals 481
E A
B
C F
D
O
P
29
that the energy dissipation capacity can be recovered for the prefabricated joint by replacing the 482
dissipater. 483
484
Fig.22. Cumulative hysteretic energy Fig.23. Equivalent viscous damping 485
coefficient 486
487
Table 3 Energy dissipation for all specimens 488
Specimen Ep,u (kN·m) ζeq,u
PJ-1 342.9 0.557
PJ-2 320.8 0.468
MJ 156.3 0.234
489
4.7 Shear deformation of joint core
As mentioned in Section 3.4, the shear drift of the joint core was gauged by extensometers set along 490
the diagonals. As shown in Fig. 24, the shear drift angle () of the joint core can be calculated as: 491 𝛾 = 𝛼1 + 𝛼2 = √𝐷2+ℎ2𝐷∙ℎ ∙ 𝛿1+𝛿22 , (3) 492
where 1 and 2 are the shear drift angles along the height and width direction of the joint core, 493
respectively; h and D are the height and width of the joint core, respectively; 1 and 2 are the 494
deformations along the diagonals. 495
0
50
100
150
200
250
300
350
400
0 30 60 90 120
Ep
(kN
•m)
(mm)
PJ-1
PJ-2
MJ
0
0.1
0.2
0.3
0.4
0.5
0.6
0 30 60 90 120
ζ ep
(mm)
PJ-1
PJ-2
MJ
30
496
Fig.24. Idealized shear deformation of the joint core 497
As shown in Fig. 25, the lateral load-shear drift angle (P-) hysteretic curves of all specimens 498
almost linearly cycled and no obvious residual deformation was observed during the initial loading 499
period. For the prefabricated joints, the development of the shear drift angle under varied loading 500
was between -0.0005 rad and 0.0005 rad, indicating that the joint core area was in the stage of 501
elastic deformation during the entire process, which will also be verified by the measured main 502
strain of the confined steel tube at the joint core in Section 4.8. Moreover, the shapes of the P- 503
curves for joints PJ-1 and PJ-2 are similar, demonstrating that the damage was controlled to take 504
place at the dissipaters of the steel hinges, which could protect the joint core. For the cast-in-place 505
monolithic joint MJ, the cracks at the joint core developed rapidly, leading to a rapid increase in 506
residual shear deformation. After the main diagonal cracks were formed, the shear drift angle 507
reached 0.006 rad. When the concrete was spalled and crushed, the joint failed due to the joint core 508
damage. 509
510
-250
-200
-150
-100
-50
0
50
100
150
200
250
-0.001-0.0005 0 0.0005 0.001
P(k
N)
(rad)
-250
-200
-150
-100
-50
0
50
100
150
200
250
-0.001-0.0005 0 0.0005 0.001
P(k
N)
(rad)
-200
-150
-100
-50
0
50
100
150
200
-0.008 -0.004 0 0.004 0.008
P(k
N)
(rad)
h
D
1
2
1
2
1
2
31
(a) PJ-1 (b) PJ-2 (c) MJ 511
Fig.25. P- hysteresis curves of all specimens 512
4.8 Strain distribution and developement
The lateral load-strain (P-ε) envelope curves for specimens were obtained by sequentially 513
connecting the extreme point of each loading level on the corresponding P-ε hysteretic curves. The 514
longitudinal strains of the longitudinal reinforcements in the beam and column (εbar), the 515
longitudinal strains of dissipaters in the steel hinge (εhinge), and the main strains of the confined tube 516
in the joint core (εtube) are presented in Figs. 26, 27, and 28, respectively. 517
As shown in Fig. 26, the longitudinal reinforcement experienced uniform strain development for 518
the prefabricated joints PJ-1 and PJ-2. The longitudinal strains did not exceed 2000 με during the 519
entire testing, that were less than the yield strains, as listed in Table 1, were 2446 με and 2527 με for 520
the beam and column reinforcements, respectively. In contrast, as shown in Fig. 27, the strains of 521
the dissipaters (εhinge) developed rapidly. After the peak, the strains of the dissipater increased 522
remarkably, the maximum value exceeding 0.02, leading to the fracture of the dissipater. These 523
results illustrate that the damage was controlled to occur at the dissipaters of the steel hinges, which 524
protected the precast beam and column from damage. Fig. 26c shows that the longitudinal strains of 525
reinforcement in the cast-in-place joint MJ were less than their yield strain; however, the 526
longitudinal reinforcements of the beam yielded before the peak point. With crack development, the 527
strain developed rapidly. 528
529
(a) PJ-1 (b) PJ-2 (c) MJ 530
0
63
125
188
250
0 1000 2000 3000 4000 5000
P (
kN
)
bar ()
Upper column
Lower column
Left beam
Right beam0
63
125
188
250
0 1000 2000 3000 4000 5000
P(k
N)
bar ()
Upper column
Lower column
Left beam
Right beam
εy
0
63
125
188
250
0 1000 2000 3000 4000 5000
P(k
N)
bar ()
Upper column
Lower column
Left beam
Right beam
εyεy
32
Fig.26. Longitudinal strain development of the longitudinal reinforcements in the beam and 531
column for all specimens 532
533
It can be seen from Fig. 27 that the strain distribution and development on the dissipater of the 534
steel hinge for prefabricated joints PJ-1 and PJ-2 were similar. For each specimen, only the strains 535
of the upper dissipater of the right beam were presented owing to symmetry. The strain gauges were 536
numbered 1 to 7 from the end connected to the joint core to the end connected to the precast beam, 537
as shown in Fig. 10b. Strain gauge No. 4 yielded first and the strain was maximum during the entire 538
process as it was installed on the weakest section. With the increase in lateral displacement, the 539
plasticity extended from the center to the ends of the dissipater, which led to excellent energy 540
dissipation. Stress strengthening allowed the joint to maintain load-bearing. In addition, strain 541
gauges No. 1 and No. 7 did not yield, limiting stress in the less ductile region near the face of the 542
column and the plastic development was concentrated in the weakened area of the dissipater. 543
Consequently, with the concept of the reduced section of the dissipater, the failure mode and 544
damage position of the proposed prefabricated joint can be controlled and the plastic hinge can be 545
outward from the beam-to-column interface. Although the energy dissipating elements are 546
symmetrically weakened, the strain measured by strain gauges No. 2 and No. 3 are larger than those 547
of strain gauges No. 5 and No. 6, as the bending moment close to the joint core is higher. 548
549
(a) PJ-1 (b) PJ-2 550
Fig.27. Longitudinal strain development of the steel hinge for prefabricated joints 551
0
63
125
188
250
0 5000 10000 15000 20000
P (
kN
)
hinge ()
1#2#3#4#5#6#7#
εy
0
63
125
188
250
0 5000 10000 15000 20000
P (
kN
)
hinge ()
1#2#3#4#5#6#7#
εy
33
Fig. 28 shows the principal stress and direction of a typical measuring point in the steel tube at the 552
joint core area for the prefabricated joints. The principal stresses calculated by the measuring stains 553
are less than the yield stress of the steel tube, indicating that the core area of the prefabricated joint 554
is confined and effectively protected by the steel tube. Thus, the seismic design concept is 555
commonly referred to as a strong connection. The direction of the principal stress in the steel tube, 556
as shown in Fig. 28 b and d, demonstrates that the steel tube was mainly subjected to a shear force. 557
558
(a) Principal stress of specimen PJ-1 (b) Principal stress direction of specimen PJ-1 559
560
(c) Principal stress of specimen PJ-2 (d) Principal stress direction of specimen PJ-2 561 Fig.28 Main stress and direction of measuring point A in the pin of prefabricated specimens 562
5. Behavior of energy-dissipating steel hinge 563
5.1 Moment and rotation curves and deformation capacity
As shown in Fig. 29, the moment-resistant (M) and the rotation () of the energy-dissipation steel 564
hinge can be calculated as: 565 𝑀 = 𝑃 ∙ 𝐻 ∙ 𝑙𝐿 (4) 566 𝜑 = 𝛥c+𝛥tℎ (5) 567
-120
-80
-40
0
40
80
120
0 5000 10000 15000 20000
s i(M
Pa)
Data acquiring times
-60
-40
-20
0
20
40
60
0 5000 10000 15000 20000
(°)
Data acquiring times
-120
-80
-40
0
40
80
120
0 5000 10000 15000
s i(M
Pa)
Data acquiring times
-60
-40
-20
0
20
40
60
0 5000 10000 15000
(°)
Data acquiring times
34
where P is the lateral load applied at the column top, H is the distance from the loading point to 568
the center of the joint core, l is the distance between the steel hinge and the center of the joint core, 569
and L is the distance between the pin support of the beam end and the center of the joint core; c 570
and t are the axial deformations of the compressed and tensioned energy dissipaters, respectively; 571
h is the height of the steel hinge section, which is defined as the distance between the centroids of 572
two energy dissipaters. 573
574
(a) Idealized internal beam-to-column joi (b) Simplified mechanical model for steel hinge 575
Fig.29 Schematic diagram of the bending moment and rotation calculation for the steel hinge 576
577
The -curves are listed in Fig. 30 for all of the specimens. As shown in Figs. 30 a and b, the 578
hysteretic curves of the energy-dissipating steel hinge of the prefabricated joint are plump in 579
shape and show sufficient flexibility. Moreover, the- curves of the left and right steel hinge 580
were almost the same for the prefabricated joint (PJ-1), as well as for PJ-2. This demonstrates that 581
the steel hinge can rotate around the pin axis. The bending moment and rotation at the points of 582
yield, limit, and failure for the steel hinge of the prefabricated joint and plastic hinge for the 583
monolithic joint are listed in Table 4. The average values were calculated as the -curves were 584
similar for the left and right hinge. Again, the characteristic values for the steel hinge in the 585
prefabricated joint under the two tests are close to each other, indicating that the function of the 586
steel hinge can be restored. Furthermore, the failure rotations were 0.048 and -0.055 for 587
t
c
Centroid
h
Section A-A
A
A
P
H
L
l
Energy-dissipating steel hinge
35
prefabricated specimen PJ-1 and 0.051 and -0.052 for PJ-2; the average value reached 0.0515, 588
which is much higher than that of the monolithic specimen MJ (the average failure rotation was 589
0.26) and demonstrated excellent rotation capacity for the steel hinge in the prefabricated joint. 590
591
(a) Specimen PJ-1 (b) Specimen PJ-2 592
593
(c) Specimen MJ 594
Fig.30 Moment-rotation hysteretic curves of replaceable energy-dissipating steel hinges 595
596
Table 4 Characteristic values of M- curves and ductility coefficients for hinges of all specimens 597
Specimen Yield point Peak point Failure point Ductility
coefficient μ
Average of μ
y (rad) My (kN·m) max (rad) Mmax (kN·m) u (rad) Mu (kN·m)
PJ-1 0.0031 176.3 0.021 253.8 0.048 215.73 15.48
16.33 -0.0032 -169.7 -0.036 -255.1 -0.055 -216.8 17.18
PJ-2 0.0033 162.1 0.031 240.9 0.051 220.1 15.45
15.37 -0.0034 -160.2 -0.033 -252.1 -0.052 -218.2 15.29
MJ 0.0028 114.6 0.0085 162.4 0.029 133.2 10.35
10.4 -0.0022 109.8 -0.0065 -187.8 -0.023 -156.6 10.45
5.2 Energy consumption ration
-300
-200
-100
0
100
200
300
-0.08 -0.04 0 0.04 0.08
M (
kN
·m)
(rad)
Left steel hinge Righe steel hinge
-300
-200
-100
0
100
200
300
-0.08 -0.04 0 0.04 0.08M
(kN
·m)
(rad)
Left steel hinge Righe steel hinge
-200
-150
-100
-50
0
50
100
150
200
-0.06 -0.04 -0.02 0 0.02 0.04 0.06
M (
kN
·m)
(rad)
Left plastic hinge Righe plastic hinge
36
The calculated values of the cumulative hysteretic energy (Ep,u) were calculated for the hinges of 598
the beam-to-column joint once the lateral load was reduced to 85% of the ultimate strength and 599
listed in Table 5. For the prefabricated beam-to-column joint, the energy dissipated by the 600
replaceable steel hinges accounted for 67.8 % during the first loading test, indicating that the 601
energy absorption was concentrated on the steel hinges. In addition, for the second loading test, 602
there was no obvious development of cracks, which led to slightly lower energy dissipation and 603
the energy consumption ratio increased to 77.0 %. The energy consumption ratio of plastic hinges 604
for a monolithic joint is 45.8 %. 605
Table 5 Energy consumption ratio for hinges of all specimens 606
Specimen Ep,u of beam-to-column joint
(kN·m)
Ep,u of hinges in joint
(kN·m)
Energy consumption ratio
PJ-1 342.9 232.4 67.8%
PJ-2 320.8 247.08 77.0%
MJ 156.3 71.56 45.8%
6. Conclusions 607
This experimental study investigated the mechanical behavior of an innovative type of 608
prefabricated beam-to-column joint subjected to cyclic loading. Based on this study, the 609
conclusions are drawn as follows: 610
(1) The connection between the precast columns and precast beams was reliable and the new 611
prefabricated beam-to-column joint exhibited good overall mechanical performance. 612
Controllable plastic hinges were formed at the ends of the precast beams because of the 613
prefabricated steel tube confined joint core and energy-dissipating steel hinge, which meet the 614
seismic design concept commonly referred to as a strong connection. The failure of the 615
specimens was concentrated on the flange of the steel hinge, primarily from the buckling of the 616
compressed flange or from fractures forming on the tensioned flange. Only minor flexural 617
37
cracks occurred near the beam and no other visible damage was observed in the prefabricated 618
beam-to-column joints. Even under repeated tests, the failure mode and damage degree of the 619
prefabricated joint can be controllable. 620
(2) The novel prefabricated beam-to-column joints exhibited excellent hysteretic behavior with 621
generally plump lateral load-displacement hysteretic curves; in contrast, the hysteretic curve of 622
the cast-in-place joint has an obvious pinching effect. Compared with the monolithic joint, the 623
load-bearing capacities of the prefabricated joint are approximately 50 % higher and the 624
deformation capacity and ductility are improved. The energy absorption capacity, in terms of 625
the equivalent hysteretic damping coefficient, was approximately two times higher than that of 626
the monolithic joint when the lateral load decreased to 85 % of the peak load. 627
(3) The initial stiffness of the monolithic joint (MJ) was slightly higher than that of the 628
prefabricated specimens; however, the stiffness degradation of specimen MJ was significant. 629
Thus, the stiffness of the monolithic and prefabricated joints was similar at 1.5 Δy. The 630
prefabricated joints exhibited no obvious strength degradation. The strength degradation 631
coefficients of the monolithic joint were approximately 0.85 to 0.90. 632
(4) In prefabricated specimen PJ-2, which only replaced the damaged dissipaters in the steel hinges 633
and was tested under repeated loading, the failure mode and P- hysteretic curve were similar to 634
those of the basic test, specimen PJ-1. However, the initial stiffness was slightly lower, 635
demonstrating that the prefabricated joint sustained a similar hysteretic performance after 636
restoration and the function of the prefabricated joints can be restored after earthquake damage. 637
(5) The - curves of the steel hinge in prefabricated joint are also in pump shape and exhibited 638
sufficient flexibility. The steel hinge can rotate around the pin axis, and the rotation at the 639
failure point reached 0.0515 rad. In the prefabricated beam-to-column joints, the energy 640
38
dissipated by the replaceable steel hinges accounted for 67.8 % and 77.0 %, respectively, 641
indicating that the energy absorption was concentrated on the steel hinges. 642
1
Acknowledgements
This study was supported by the National Natural Science Foundation of China (Grant No.
51578152, 51878174) and the Natural Science Foundation of Fujian Province (2020J01887). The
financial support is highly appreciated.
Declaration of Competing Interest
The authors declare that they have no known competing financial interests or personal relationships
that could have appeared to influence the work reported in this paper.
Data Availability Statement
The data that support the findings of this study are available from the corresponding author upon
reasonable request.
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Declaration of interests
☒ The authors declare that they have no known competing financial interests or personal relationships
that could have appeared to influence the work reported in this paper.
☐The authors declare the following financial interests/personal relationships which may be considered as
potential competing interests: