The Seismic Performance of Self-Centering Ribbed Floor Flat-Beam Frame Joints

Abstract

To achieve rapid post-earthquake repair of self-centering ribbed floor flat-beam frame structures, a ductile hybrid joint consisting of dog-bone-shaped, weakened, energy-dissipating steel bars connected to the upper and lower column sections through high-strength threads is proposed based on the damage control design concept. By moving the ductile energy-dissipating zone out to the locally weakened section of the energy-dissipating steel bars and the locally unbonded prestressed steel bars in the core area, the residual deformation was limited and the seismic performance improved. Based on the working principle of hybrid joints, low cycle loading tests were conducted on two joint specimens to analyze the influence of lateral prestress on the seismic performance of the hybrid joints. Numerical modeling methods were used to compare the position of the energy-dissipating steel bars in the composite layer and the friction performance of the joints. The research results indicated that the hybrid joint had stable load bearing, deformation, and energy dissipation capabilities, with damage being primarily concentrated in the energy-dissipating steel bars. Even at an inter-story displacement angle of 5.5%, the upper and lower column segments remained elastic. After unloading, the connection seam at the joint was closed, and the self-centering performance was good. When the inter-story displacement angle reached 5.5%, the lateral prestress increased from 150 kN to 250 kN, the ultimate bearing capacity of the joint increased by 16.3%, and the cumulative energy consumption increased by 30.0%. The influence of the friction coefficient of the joint surface on the structural performance was set at a threshold of 0.7. When it was less than the threshold, the ultimate bearing capacity and initial stiffness of the joint increased with the increase in the friction coefficient. After reaching the threshold, the increase in the ultimate bearing capacity of the joint slowed down, and the rate of stiffness degradation gradually accelerated. This joint showed excellent seismic performance and can thus achieve post-earthquake repair of structures.

1. Introduction

The self-centering ribbed-slab beam frame structure (IMS system) was proposed in 1956 by Zezely, a Professor from Yugoslav Serbian Materials Research Institute. It is a structure composed of a prefabricated slab and column prestressed by the post-tensioning method. Scholars have conducted extensive research on prestressed, prefabricated friction joints and self-centering structures. However, research on the friction torque mechanism and energy dissipation contribution of the IMS system under seismic load is still lacking. Therefore, based on the working principle of hybrid joints, this paper proposed new IMS joints that can achieve rapid post-earthquake repair of structures.

The National Institutes of Standards and Technology (NIST) [1,2,3,4,5] research project and the Precast Seismic Structural Systems (PRESSS) program in the United States have proposed the use of prestressing technology to control the residual deformation of prefabricated concrete structures while reducing damage to structural components by dissipating energy through joint energy dissipation systems. Morgen [6] designed and conducted experiments on setting friction-damping energy dissipation devices at both ends of prestressed, prefabricated beam–column joints. The results indicated that prestressed, prefabricated friction joints equipped with friction dampers had good energy dissipation performance. Huang Linjie [7,8] completed 14 sets of prestressed concrete frame joint tests with variable friction dampers (VFDs) placed in the joint area. The results showed that this type of joint had good load-bearing capacity and ductility while effectively solving the problems of the high-order modal effects of self-centering structures, the reduced stiffness of joints after pressure reduction, and excessive inter-story lateral displacement. The problem of the difficult installation of energy-consuming steel bars at the bottom of prefabricated beams with prestressed prefabricated joints remains to be solved.

The recentering materials in components consist of prestressed tendons and steel strands [9], fiber-reinforced composite material bars with shape memory alloys, springs, etc. [10,11,12]. To promote the application of self-centering structural components in engineering, Yan et al. [13] proposed a self-centering joint between CFDST columns and RC beams. Analytical studies have shown that increasing the friction can improve joint energy dissipation, but it will increase the residual deformation of the joint. An experimental study looked at using martensitic NiTi bars in a beam–column connection [14] in which SMA screws were subjected to 0.5% prestrain to ensure the initial stiffness of the beam–column joint, and the test results show that the joint can still maintain its self-centering ability when it reaches a 5% corner angle. Self-centering reinforced concrete joints based on SMAs were proposed by Youssef et al. [15]. The SMA rod and steel bar were coupled with a steel sleeve fastened by a single cylindrical bolt, and the experimental joint demonstrated good deformation recovery.

Wang Xianming [16] completed a quasi-static test of a prestressed IMS slab column interlayer model with four columns and two panels. The results showed that the frictional bending moment accounted for approximately 45% of the ultimate bending capacity of the joint. Aiming to solve the problems of poor energy dissipation in the IMS structure, Hongyu Chen [17,18] determined the optimal tooth groove structure for the prefabricated components of an extra-large column grid IMS and calculated the friction coefficient of different joint structures through experiments, determining the analysis parameters for numerical simulation of the joint. Experimental studies have shown that the seismic performance of hybrid joints can be comparable to or even exceed that of cast-in-place joints, and they have the characteristics of self-centering and low loss.

In terms of self-centering beam–column joints, Guo Haishan [19,20] proposed an asymmetric reinforcement structure of precast, prestressed, efficiently fabricated frame (PPEFF) joints and conducted joint and overall frame tests. The research showed that the PPEFF system had a fast construction speed, low labor consumption, and excellent seismic and continuous collapse resistance performance. Jule Zheng [21] compared the seismic resilience of ordinary RC structures with that of flag-shaped hysteretic behavior and proved that in their proposed self-centering, prestressed concrete frame with a generalized flag-shaped hysteretic, the damage mainly came from the damage of acceleration-sensitive non-structural components in the upper floor. Jiang et al. [22] put forward a self-centering prestressed prefabricated steel beam–column joint with a weakened dog-bone-shaped flange cover plate, and carried out finite element simulation and experimental studies. The results show that the joints have good bearing capacity, seismic performance, and self-centering capacity after an earthquake.

Based on the above research on the self-centering ability and energy dissipation ability of structures, this study proposes a ductile hybrid joint that is connected to the upper and lower column segments by high-strength threaded connections using weakened dog-bone-shaped energy-dissipating steel bars. Considering the connection effect of the slab haunching area on two-directional composite beams, further research is conducted through experiments and numerical simulation to investigate the influence of frictional torque at the end of the composite beam on the bearing capacity, stiffness degradation, energy dissipation performance, ductility characteristics, and failure modes of the joints.

2. Overview of Hybrid Joints

A hybrid joint with replaceable energy-consuming steel bars was constructed, as shown in Figure 1. The concrete in the core area of the joint was in a state of triaxial compression. The difference between the lower prefabricated and upper cast-in-place composite beams of the prestressed, prefabricated frame structures and the upper cast-in-place composite beams is that the IMS system composite beams are “sandwich”-shaped (prefabricated at both ends and cast-in-place in the middle), and the vertical connection surface of the “sandwich”-type composite beam is more conducive to bearing the bending moment at the beam end than the horizontal connection surface of the upper and lower composite beams.

As shown in Figure 2a, for the IMS system of rectangular columns, the bending moment on one surface becomes the torque on another surface, and the torque in turn becomes the bending moment on another surface, which is the torsional moment generated by the frictional shear force. This torque is called the frictional torque at the end of the composite beam. The unique frictional torque at the end of the laminated beam in the IMS system joint can achieve relative sliding friction energy dissipation through the plate column interface under large deformation, which has the same effect as the energy-dissipation devices added outside the joint. The bending friction torque under the ultimate load is as follows:

$$M_{TU}=\frac{\mu N_{lat}b(6h-2b-3x_c)}{6h}$$

(1)

where $\mu$ is the coefficient of friction, $b$ is the width of the side seam surface, $N_{lat}$ is the comprehensive axial force of lateral prestressing, $h$ is the height of the plate rib, and $x_c$ is the height of the concrete compression zone [23].

The spatial deformation and force principle of the joint are shown in Figure 2b. Under strong earthquakes, the relative rotation between prefabricated columns and prefabricated slabs causes close contact on one side of the prefabricated slab and column, forming a rotation center. The open channel on the closer side undergoes slight compression deformation, while the open channel on the farther side undergoes significant elongation deformation, causing plastic damage to primarily concentrate on the contact section between the open channel and the prefabricated column.

3. Experimental Overview

3.1. Test Piece Design

To study the seismic performance and functional recoverability of hybrid joints, two specimens with the same overall geometric dimensions were designed in the experiment, numbered A1 and A2. The overall dimensions of the specimens are shown in Figure 3a. The total height of the prefabricated column was 1600 mm, with a cross-sectional size of 360 mm × 360 mm. The concrete strength of the open channel and the composite layer was C40. The thickness of the protective layer of the specimen was 15 mm, the width of the open channel was 100 mm, the rib height of the prefabricated floor slab was 150 mm, and the thickness of the composite floor slab was 120 mm. The magnitude of the lateral prestress in the experimental design was used as a variable to study the influence of the frictional torque at the end of the composite beam on the seismic performance of the joint. A total initial prestress of 400 kN was applied to the longitudinal open-channel prestressed reinforcement. Design specimens A1 and A2 with lateral prestress of 150 kN and 250 kN respectively, with a protective layer thickness of 15 mm. Holes were conserved inside the prefabricated column into which steel bar connectors were embedded. The research on the bearing capacity failure and failure mechanism of the cast-in-place beam slab column joints under earthquake action is relatively mature, so no comparative tests were conducted on the cast-in-place beam slab column joints.

The structure of the specimen joint is shown in Figure 3b. The longitudinal bars connecting the upper and lower portions of the composite beam to the column were locally weakened and treated without bonding. They were arranged in an open channel, and the longitudinal bars were connected to the pre-embedded steel bars in the column through a steel thread connector. The prestressed bars near the core area of the joint were changed from originally bonded to locally unbound. The prestressed steel strands were wrapped with double-layer polyethylene tape for local unbound treatment, and the length of the unbonded section was 500 mm. The length of the weakening section of the energy-consuming steel bar was 120 mm, and the proportion of the section weakening was 15%. The prefabricated components and prestressed steel bars were made of 1 × 7 Φ S15.2 steel strand with a steel strand strength of fptk = 1928 MPa and fpy = 1786 Mpa and designed according to current specifications [12]. The size of the structure and reinforcement are shown in Figure 4.

3.2. Material Properties

According to GB/T 228.1-2010 [24], “Tensile testing of metallic materials—Part 1: Room temperature test method,” samples were collected from the internal steel bars of various key components (all HRB400 grade), and material tensile tests were conducted. The mechanical performance indicators of the steel bars with different diameters are shown in

Table 1. Mechanical properties of steel.

Steel Type	Diameter (mm)	Strength/Mpa		Elastic Modulus/(Mpa)	Elongation Agt/%
		Yielding	Ultimate
Reinforced bar	16	482	657	2.17 × 105	20.1
	12	441	603	2.05 × 105	22.6
	8	428	584	2.11 × 105	23.7
	6	406	552	2.05 × 105	24.3
Strands	1 × 7 Φ S15.2	1786	1928	1.98 × 105	6.5

Note: The mechanical property data are the averages of the test results of 3 specimens.

.

The design strength grade of the precast concrete components was C40. At the same time each specimen was poured, 150 mm × 150 mm × 150 mm cubic test blocks were constructed and cured under the same conditions as the components. The average compressive strength of the concrete cube measured during specimen loading was 41.0 Mpa [25]. To improve the early compressive strength and final flexural strength of the grouting material, a high-strength, non-shrinkage grouting material mixed with steel fibers was selected. The material properties are shown in

Table 2. Test mix proportion parameters and results.

Water Cement Ratio	Volume Content/%	Compressive Strength/MPa			Flexural Strength/MPa			Fluidity /mm
		1 d	3 d	28 d	1 d	3 d	28 d	Initial Value	30 min
0.145	1.5	65.0	72.7	80.5	8.4	11.8	10.9	305	285

.

3.3. Experimental Loading and Measurement Plan

3.3.1. Loading Device

The repeated low cycle loading test was conducted at the Engineering Structure Center of Hebei Agricultural University. The loading device is shown in Figure 5a, and it consisted of one 1000 kN vertical hydraulic jack, two 500 kN bidirectional CNC hydraulic jacks with ball joints, one set of ball joint column caps with horizontal pull rods, and a reaction frame. The hydraulic jack was connected to the loading point at the plate end using anchoring screws, steel clamps, and vertical hydraulic jacks to apply axial pressure to the top of the column, and these could freely slide along the horizontal direction with the specimen. To prevent out-of-plane overturning, rigid triangular supports were installed at a distance of 20 mm from both sides of the specimen and fixed to the rigid ground by anchor bolts.

3.3.2. Measurement Plan

The arrangement of the displacement gauges is shown in Figure 5b. Displacement gauges D1—D8 were used to measure the displacement that occurs during the loading of the specimen. Displacement gauges D1 and D2 measured the vertical displacement at the loading point of the plate end, while displacement gauges D3 and D4 were used to monitor the vertical and horizontal displacement of the column respectively during the loading process. Displacement gauges D5—D8 were placed at the connection between the prefabricated column and the composite beam to measure the relative rotation of the joint. Strain gauges were symmetrically arranged on both sides of the open channel at the bottom of the composite beam at the core area of the joint to monitor the strain changes of the cross-section at the maximum bending moment. In addition, areas with complex local strain distributions, such as the haunching area of the prefabricated panels and the upper portion of the prefabricated panels, were sprayed with white latex paint to monitor plastic development. During the test, the measured data of the load sensor and displacement meter were collected by a Donghua dynamic data acquisition instrument (model number: DH5902), and the load and displacement changes at the beam end were observed in real time during the test. The strain gauge data were automatically collected after the collection frequency was set using the static data acquisition instrument (model number: DH3816).

3.3.3. Loading System

The experiment used quasi-static cyclic loading. First, a vertical axial force was applied at the top of the column according to the nominal axial compression ratio (0.32) designed, and this was kept constant. We then used vertical displacement control loading at the plate end. The vertical displacement loading system referred to the JGJ/T101-2015 “Code for Seismic Testing of Buildings” standard [26], as shown in Figure 6. Before formal loading, preloading was used to check whether the specimen connection was secure and whether the instrument was working properly. Loading was conducted step by step to evaluate the remaining bearing capacity of the IMS joints after a major earthquake and to ensure loading safety. The final inter-story displacement angle was loaded to 5.5%. The vertical bidirectional jack at the slab end was used to control the displacement of the slab end and the antisymmetric applied load. The vertical actuator was loaded through displacement control, and the ratio of the vertical displacement of the loading point to the horizontal distance from the outer edge of the column is defined as the drift ratio θ. Specifically, when θ is 1/2000, cycle once; when θ is 1/1000, 1/800, 1/550, 1/400, 1/300, 1/200, 1/100, 1/67, 1/50, 1/40, 1/30, 1/25, 1/20, or 1/18, cycle 3 times. During the loading process, it was specified that the vertical displacement of the bidirectional tension and compression jack was positive when extrapolated and negative when pulled back.

4. Test Result and Discussions

4.1. Experimental Phenomena and Failure Modes

Cracks are local fracture phenomena in concrete structures caused by internal stress or external factors. Cracks can be divided into surface cracks and internal cracks. Surface cracks can be directly observed, while internal cracks may require non-destructive testing techniques and other means for detection. The cracks mentioned in this article are all surface cracks observed during the experimental loading process. The test phenomena of the A1 and A2 test specimens were similar. Significant cracks appeared at the joint between the plate and the column as the loading progressed in test specimen A1, and eventually, there was significant opening and closing. When the inter-story displacement angle reached 1%, cracks appeared on the surface of the plate above the lateral open channel and expanded toward the direction of the prefabricated column. Vertical cracks appeared in the haunching area at the bottom of the prefabricated plate. The damage severity is used to describe the overall performance degradation of concrete structures under the influence of cracks and other damage factors. This paper adopts a damage degree evaluation based on crack parameters. By counting the number, length, width, and depth of cracks, combined with the size and stress situation of the concrete structure, the damage severity is calculated. When the inter-story displacement angle reached 2.5%, the concrete in the haunching area at the bottom of the prefabricated plate was crushed and peeled off. Figure 7 shows that as the interlayer displacement angle increased by 5%, the maximum opening gap between the plate and column reached 18 mm. Due to the repeated tensile and compressive buckling of the energy-consuming steel bars, the concrete at the bottom of the exposed channel slab swelled, and the concrete protective layer at the root of the composite beam collapsed and peeled off. After loading stopped, there was significant residual deformation at the joint. Throughout the entire loading process, the prestressed steel bars were basically in an elastic state, and the lateral bending friction torque provided by the lateral prestressed steel bars always provided effective constraints.

The longitudinal prestress of specimen A2 was the same as that of specimen A1, but the lateral prestress was different. During the early stage of loading, the test phenomenon was similar to that of specimen A1. After joint cracking, the energy-consuming steel bars deformed. However, as the inter-story displacement angle increased, the column corner was crushed and peeled off at the intersection of the plate and column. When the inter-story displacement angle reached 5%, the opening width at the joint reached 16 mm, and the concrete at the bottom of the laminated beam that connected to the column was crushed and peeled off. The horizontal cracks on the side of the haunching area at the bottom of the slab continued to expand and extend. When loaded to the final displacement angle, the specimen’s floor experienced significant out-of-plane displacement, causing tearing between the prefabricated panel and the laminated layer. When the test loading was completed, the joint still maintained good shear resistance and self-centering ability.

4.2. Hysteresis Curves and Skeleton Curves

The hysteresis curves of the plate end load interlayer displacement angle for each specimen are shown in Figure 8. It can be seen that the hysteresis performances of the joints in the positive and negative directions were different due to the influence of the floor slab. Due to the joints’ characteristics of “small deformation rigid consolidation and large deformation flexible energy dissipation”, they resisted the shear force at the beam end and consumed seismic energy, resulting in a fuller hysteresis curve after increasing the lateral constraint load. During the initial stage of loading, the hysteresis curve showed a linear variation. As the inter-story displacement angle increased, the energy-consuming steel bars gradually yielded. As the number of cyclic loadings increased, the concrete at the positive end of the slab cracked, and the joints cracked into plastic hinges; the hysteresis curve showed obvious bending, gradually transitioning to a bow shape, with significant pinching phenomenon. After the specimen reached its peak load, the bonding force between the energy-consuming steel bars and concrete gradually failed, and the overall stiffness of the structure decreased more significantly. The hysteresis curve becomes fuller, and the residual deformation rapidly increases after unloading. A comparison showed that an improvement in the lateral prestress significantly improved the seismic performance of the joints under large earthquakes. This was primarily because the lateral prestress was directly proportional to the lateral friction torque of the composite beam column. When the bidirectional friction joint bent and deformed in one direction, the prestress in the other direction produced compression, embedding, and constraint balance effects, thereby improving the energy dissipation performance of the joint.

Figure 9 shows the skeleton curves of each specimen. It can be seen that the development trends of the skeleton curves of the two specimens were basically consistent. After the yielding of the energy-consuming steel bars, each specimen entered the plastic strengthening stage, and the load on each specimen increased. After reaching the peak load, the bond strength between the prestressed steel bars and the concrete was gradually destroyed, and a decreasing section quickly appeared. After increasing the lateral prestress of the specimen by 100 kN, the positive peak load increased by 16.3% and the negative peak load increased by 17.8%. In addition, it was calculated that the positive displacement ductility coefficient and negative displacement ductility coefficient of specimen A2 were increased by 18.9% and 26.3%, respectively, compared to specimen A1, indicating that an increase in the lateral prestress improved the ductility of the joint and significantly increased its seismic toughness.

4.3. Accumulated Energy Consumption

To quantitatively analyze the energy consumption capacity of the specimen, the cumulative energy consumption [19] (Ed) of the specimen was used as the measurement indicator. The larger the Ed value, the better the energy dissipation capacity of the specimen. Based on the hysteresis curve of the specimen in Figure 8, the area of each half-cycle hysteresis loop was calculated and accumulated sequentially to obtain the cumulative energy consumption change curve of each specimen, as shown in Figure 10. As the number of loading cycles increased, the cumulative energy consumption of each specimen continuously increased. The cumulative energy consumption curve of specimen A2 was consistent with the basic trend of specimen A1, showing a different energy consumption performance after a displacement angle of 1.5%. When the displacement angle was 5.5%, the cumulative energy consumption of specimen A2 was 30.0% higher than that of specimen A1. The experiment showed that increasing the frictional torque at the end of the composite beam enhanced the constraint effect of joint torsion. Due to the plastic failure of the plate bottom haunching area and the plate column joint surface, most of the energy was absorbed. In summary, under large deformation conditions, the frictional torque at the end of the composite beam made a significant contribution to the energy dissipation of the joint.

5. Finite Element Analysis

5.1. Model Establishment

As shown in Figure 11, ABAQUS 6.14 software was used to model the experimental joints in a three-dimensional solid manner. The geometric dimensions, boundary conditions, and loading methods of the model were consistent with those of the experiment. The separated modeling method was used, and steel that used a three-dimensional truss linear element, T3D2, concrete, and steel plates all used solid elements. Considering the effects of large deformation and contact in the simulation, C3D8R elements were used for the edge beams, columns, and exposed channel concrete. Due to the large number of contacts and elements in this model, to save calculation time and improve calculation accuracy, local refinement was carried out on the transverse open-channel concrete mesh. In the FEA, the dimensions of the channel slab and column units are 50 mm, the locally increased dimensions of the longitudinal open groove are 35, the dimensions of the steel plate unit are 20 mm, and the dimensions of the steel bar and steel strand unit are 20 mm. Linear non-coordinated-mode elements, C3D8I, were used for the joint mortar and prefabricated floor slabs to suppress the hourglass effect during numerical calculations. To more effectively simulate the pinching effect of the hysteresis curves, the PQ fiber [27,28,29,30] user material subroutine developed by Tsinghua University was used. The concrete plastic damage (CDP) model considers the constraint of hoop reinforcement on concrete and uses different uniaxial compressive stress–strain relationships of concrete to describe the components in the model. For the composite layer and post-poured concrete in the open channel, the plain concrete model provided in [31,32,33] was used. For the concrete in prefabricated columns and slabs, the model provided by Légeron [34,35] was used.

5.2. Finite Element Results and Analysis

A comparison of the hysteresis curve test and simulation results of the repeated low cycle loading of joints is shown in Figure 12. It can be seen that the overall trend of the simulation results was the same as that of the test hysteresis curve. The relative error between the ultimate test load and the finite element simulation load was basically within 15%. The unloading stiffness in the simulation matched well with the test’s hysteresis curve. Due to the idealization of the boundary conditions and the connection between the test components in the finite element model, there was a certain degree of separation between the finite element simulation curve and the test curve. However, overall, the finite element results were in good agreement with the test results, indicating that the established finite element model better simulated the stress characteristics of joints in IMS under repeated loads.

Figure 13a,b show the stress cloud maps of the steel reinforcement skeleton and concrete at a loading displacement angle of 1/20 for joint A1. It can be seen that the stress on the upper left end and lower right end of the composite beam at the joint was relatively high, primarily due to the large bending moment at that location. The bottom energy-dissipating steel bars increased the ductility of the joint under repeated action during the yield stage, which was consistent with the deformation of the joint area and the strain of the steel bars during the experiment. According to the Mises stress cloud map of the concrete cross-section in the core area, there was a significant stress concentration in the column corner area and the haunching area at the bottom of the slab. This result was in good agreement with the failure mode observed during the loading process.

5.3. Parametric Analysis

Figure 14 shows that due to the prestressed assembly ribbed plate column structure, the prefabricated plate corners had grooves and were connected to the columns on both sides, which was different from the traditional single-sided flat connections of prefabricated beams and columns. The stress on the two faces of the plate corners was interrelated. For the rectangular-column IMS joints, the bending moment on one face will become the torque on the other face, and the torque will become the bending moment on the other face. This is the torsional moment generated by frictional shear force that is called the frictional torque at the end of the composite beam. The unique frictional torque at the end of the laminated beam in the IMS system joint achieved relative sliding friction energy dissipation through the plate column interface under large deformation, and this had the same effect as the energy dissipation devices added outside the joint. To investigate the influence of frictional torque at the end of the composite beams on the seismic performance of the prestressed assembled ribbed plate column joints, the numerical model of specimen A1 was used as an example. In addition, the friction coefficient of the plate column contact surface and the thickness of the post-poured composite layer were selected as the primary influencing factors for parameter analysis.

5.3.1. Friction Coefficient

The friction coefficient of the plate column contact surface is an important parameter that affects the frictional torque at the end of the laminated beam in IMS joints. Referring to the recommended values in the specifications and considering unfavorable factors such as poor construction quality of the contact surface, the friction coefficients were set at 0.2, 0.4, 0.6, 0.7, 0.85, and 0.9 for the parameter analysis. The load displacement angle skeleton curves under different friction coefficients are shown in Figure 15. It can be seen that the friction coefficient of the plate column joint surface had a significant impact on the bearing capacity, stiffness, and ductility of the joint. The larger the friction coefficient of the joint surface between the plate and column, the greater the peak load of the joint. During the early stage of loading, the joint was in the elastic stage, and the initial stiffness curve trend was basically consistent. After the specimen yields, the maximum bearing capacity of the joint increased with an increase in the friction coefficient, but the increase was no longer significant after reaching 0.7. When the friction coefficient of the specimen exceeded the ultimate load, the skeleton curve was relatively stable. The joint had good ductility, but the bearing capacity was not high. Therefore, in practical engineering, structural measures should be taken to ensure that the friction coefficient of the plate column joint surface is greater than 0.7 to ensure that the joint has high bearing capacity and ductility.

5.3.2. Laminated Layer

Due to the non-prestressed state of the composite layer and the post-poured concrete in the exposed channel, when the precast component concrete was damaged, the post-poured concrete in the composite layer and the exposed channel did not reach its ultimate compressive strain. The setting of the composite layer helped to improve the ductility of the joint and delay a decrease in the joint’s bearing capacity. The thicknesses of the composite layer were set to 50, 75, and 100 mm for the parameter analysis. The load displacement angle skeleton curves for the different thicknesses of the composite layer are shown in Figure 16. It can be seen that the composite layer had a significant impact on the bearing capacity and ductility of the joint after yielding. As the thickness of the composite layer increased, the bearing capacity of the joint increased at the same displacement angle. The peak load of the specimen with a composite layer thickness of 50 mm was approximately 21.32% greater than that of the specimen without a composite layer. The peak load of the specimens with a composite layer thickness of 75 mm and 100 mm was 7.6% and 15.38% greater than that of the specimen with a composite layer thickness of 50 mm, respectively. The bearing capacity of the specimen without a composite layer decreased significantly after the peak load, and the ductility and integrity were poor. Therefore, to ensure high strength and ductility in IMS joints, the thickness of the cast-in-place layer of the composite floor slab should be greater than 75 mm and a certain amount of composite floor slab reinforcement should be configured.

The prestressed prefabricated system should not only have a certain seismic bearing capacity but also have good deformation and consumption capacity to ensure that the structure can resist collapse. Controlling the axial compression ratio plays a crucial role in controlling the ductility of the structure. During the IMS system specimen testing process, the axial compression ratio was 0.32. Numerical simulation software was used to calculate the skeleton curves of the specimens at axial compression ratios of 0.3, 0.4, 0.5, 0.6, 0.7, and 0.8, as shown in Figure 17. From the figure, it can be seen that in the initial loading stage, the skeleton curve of the specimen does not change with different axial compression ratios. When the specimen enters the elastic–plastic stage, the axial compression ratio has a certain degree of influence on the ultimate bearing capacity and ductility of the specimen. As the axial compression ratio increases from 0.3 to 0.6, the ultimate bearing capacity of the specimen’s skeleton curve also increases. When the axial compression ratio increases to 0.7 and 0.8, the ultimate bearing capacity of the specimen’s skeleton curve does not change significantly. However, when the axial compression ratio is 0.7 or 0.8, the peak load of the specimen decreases significantly in the later stage. Due to the structure entering the elastic–plastic stage when the axial compression ratio is high, the damage to the column increases, leading to local failure and reduced ductility. With the increase in the axial compression ratio, the peak load of the joint increases, and the magnitude of the strength decreases after the peak load also increases, resulting in a decrease in ductility. Therefore, in the seismic design of the IMS structural system, the limit value of the column’s axial compression ratio should be taken according to the requirements of specifications.

6. Conclusions

(1) The new type of IMS hybrid joint proposed in this study has good load-bearing and energy-dissipation capabilities, as well as strong seismic performance. It can effectively bear loads through the collaborative working mechanism between the prefabricated components. At the same time, the dual energy-dissipation mechanism combining the plastic deformation of steel bars in open channels and the frictional torque between plates and columns is fully utilized to dissipate energy. Even at an inter-story displacement angle of 5.5%, the upper and lower column segments remained elastic. After unloading, the connection seam at the joint was closed, and the self-centering performance was good. The lateral prestress increased from 150 kN to 250 kN, and the ultimate bearing capacity of the joint increased by 16.3%. When the loading displacement angle was 5.5%, the cumulative energy consumption was 30.0% higher.

(2) Under repeated low cycle loads, the IMS hybrid joint can control the plastic damage sustained during an earthquake on easily replaceable energy-consuming steel bars, while the unbonded prestressed steel bars are basically in an elastic working state. After the earthquake, only the energy-consuming steel bars in the exposed groove need to be replaced, achieving rapid repair of the structure’s post-earthquake function. Its excellent exposed groove construction provides the possibility for multiple repairs. Setting a composite layer can not only enhance the overall integrity and reliability of a structure but can also help improve joint ductility and delay a decrease in the joint bearing capacity.

(3) The friction coefficient threshold of the joint surface is 0.7, which can impact the structural performance. When it was less than the threshold, the ultimate bearing capacity and initial stiffness of the joint increased with an increase in the friction coefficient. After reaching the threshold, the increase in the ultimate bearing capacity of the joint slowed down, and the rate of stiffness degradation gradually accelerated. This joint had excellent seismic performance and also achieved post-earthquake repair of the structure.

Future research should consider the contribution of bracket construction to the shear-bearing capacity and safety of joints. This will very much be the key component in future attempts to improve the assembly and construction efficiency of the structural system and reduce the use of temporary supports. Further studies should focus on the seismic performance and bearing capacity calculation methods of IMS structures with brackets. By establishing a three-dimensional refined model of the IMS shear wall structural system and analyzing the post-earthquake damage of the structure under rare seismic intensities, the application prospects of the new IMS structural system in high-strength areas and high-rise buildings can be comprehensively evaluated.