Experimental Investigation on Static Performance of Novel Precast Concrete Composite Slab–Composite Shear Wall Connections

Abstract

The connection zones between precast concrete composite slabs and composite walls commonly experience severe reinforcement conflicts due to protruding rebars, significantly reducing construction efficiency. To address this, a novel slotted concrete composite slab–composite shear wall (SCS-CW) connection without protruding rebars is proposed in this study. In this novel connection, rectangular slots are introduced at the ends of the precast slabs, and lap-spliced reinforcement is placed within the slots to enable force transfer across the joint region. To investigate the static performance of SCS-CW connections, four groups of connection specimens were designed and fabricated. Using the structural detailing of the connection zone as the variable parameter, the mechanical performance of each specimen group was analyzed. The results show that the specimens demonstrated bending failure behavior. The key failure modes were yielding of the longitudinal reinforcement in the post-cast layer, yielding of the lap-spliced reinforcement, and concrete crushing at the precast slab ends within the plastic hinge zone. Compared to composite slab–composite wall connections with protruding rebars, the SCS-CW connections demonstrated superior ductility and a higher load-carrying capacity, satisfying the design requirements. Additionally, it was revealed that the anchorage length of lap-spliced reinforcement significantly affected the ultimate load-carrying capacity and ductility of SCS-CW connections, thus highlighting anchorage length as a critical design parameter for these connections. This study also presents methods for calculating the flexural bearing capacity and flexural stiffness of SCS-CW connections. Finally, finite element modeling was conducted on the connections to further investigate the influences of the lap-spliced reinforcement quantity, diameter, and anchorage length on the mechanical performance of the connections, and corresponding design recommendations are provided.

1. Introduction

Prefabricated construction refers to a building method in which various structural components are manufactured in advance at a factory, transported to the construction site, and then assembled and connected to form an integrated building structure [1,2]. Prefabricated construction, with its wide-ranging benefits, is leading the shift and advancement of the construction sector, surpassing traditional cast-in-place approaches in areas such as construction efficiency, stability, environmental friendliness, and labor costs [3,4,5]. Its features of standardized production and modular assembly significantly enhance project controllability and have become an important direction for innovation and development in the industry. Prefabricated construction adopts modular design, offering a high degree of flexibility and scalability. Modules with different functions can be combined and adjusted according to specific requirements, meeting diverse architectural needs. As a result, prefabricated construction has been widely applied in various types of buildings, including high-rise structures, residential buildings, office buildings, and public facilities [6].

Concrete composite slabs and composite shear walls are typical components in prefabricated construction. The construction method for concrete composite slabs features the combination of prefabricated elements with a cast-in-place topping layer, forming an integrated load-bearing system through layered casting. This approach effectively merges the precision of factory prefabrication with the adaptability of on-site construction, making it widely applicable in industrial plants, residential buildings, and other engineering fields [7,8,9,10]. During the construction stage, the precast slab can serve as a substitute for traditional formwork, facilitating the casting of the topping layer; once the topping material reaches the required strength, it is integrated with the precast slab to form a monolithic structure that bears loads cooperatively [11,12]. In prefabricated construction, the precast concrete shear wall structure is the primary structural form. Precast wall panels are connected by reinforcement cages to form a stable formwork system, and an integrated load-bearing member is created by casting concrete into the core zone on site [13,14,15,16]. This structural system combines the high precision of factory fabrication with the continuity advantages of cast-in-place structures, achieving effective cooperation between prefabricated components and post-cast layers [17,18]. Composite shear wall structures are categorized as single-sided or double-sided types [19,20], as shown in Figure 1. During the construction phase, the precast concrete walls serve as formwork, and in the service phase, both sides of the double-sided composite shear wall participate in structural load bearing. In contrast, in single-sided composite shear walls, the external precast panel does not contribute to structural load resistance [21].

Connections represent the most critical zones within a structural system [22], and numerous studies have focused on connection methods for precast concrete joints. Many researchers have investigated the use of post-tensioned prestressing tendons passing through the joint region to enhance the overall performance of precast connections [23,24,25,26,27]. Findings have shown that this type of connection can significantly improve the stiffness and ductility of the joint. Hybrid joint systems have also emerged as a key area of interest in the study of precast concrete connections. In these systems, steel components are embedded within the precast elements, and load transfer between concrete components is primarily achieved through steel plates, bolts, or similar mechanical interfaces. Eom et al. [28] developed a novel steel bracket system installed at precast beam–column joints, which enabled effective moment transfer without damage to the connection. Wu [29] and Eray Baran [30] proposed connection methods using angle steel and high-strength bolts to link precast beams to concrete columns. Zhu et al. [31] introduced a new type of concrete-filled steel tube beam–column joint and focused on analyzing its failure mechanisms and energy dissipation capacity. As for the mechanical performance of wall–slab connections, S. R. Salim et al. [32] introduced additional stirrups in the wall–slab connection region, and test results revealed that increasing the number of stirrups in the connections led to improved ductility and higher energy absorption capacity. D. Zenunovic et al. [33] investigated the connections linking precast slabs to monolithic walls and monolithic slabs to walls through experiments. The results indicated that the designed precast connections achieved a load-carrying capacity comparable to that of monolithic connections, although further hysteretic tests were recommended for deeper investigation. Ding et al. [34] developed a wall–slab structure based on UHPC sandwich panels. A. Chalot et al. [35] applied fiber-reinforced polymer (FRP) to wall–slab connections and conducted hysteretic tests, demonstrating that FRP reinforcement significantly improved both the strength and energy dissipation capacity of the connections.

For the connection methods of precast concrete wall panels, the traditional approach commonly adopts the use of protruding reinforcement at the ends of precast components. Load transfer is achieved by anchoring the protruding rebars into the post-cast layer, as illustrated in Figure 2. However, this construction method faces the following challenges: first, the positioning requirements for the protruding reinforcement demand high precision during the fabrication of precast components; second, during assembly, the protruding rebars from different components are highly prone to conflicts with each other and with the surrounding concrete, causing significant inconvenience during construction and reducing construction efficiency, which is inconsistent with the design philosophy of prefabricated construction [36].

To address these issues, researchers have carried out a series of studies focusing on the connection configurations of precast concrete structures. Jiang et al. [36] proposed a connection method in which a slot is formed at the end of the precast slab and anchorage reinforcement is placed within the slot. Subsequently, Yang et al. [37] investigated the bearing capacity of slotted tightly jointed concrete composite slabs and found that enhancing the end anchorage of additional reinforcement could further improve the bearing capacity of the precast specimens. Nie et al. [38] conducted mechanical performance tests on bidirectionally loaded slotted concrete composite slabs, and the results indicated that this connection could achieve effective force transfer across the joints and exhibited mechanical behavior similar to cast-in-place concrete slabs. Huang et al. [39] introduced a method where slots were formed in bridge decks of a certain thickness to replace protruding rebars, with reinforcement placed within the slots to resist shear forces, thus achieving shear connection between steel beams and precast bridge decks. While existing lap-spliced and slotted designs are widely used for connecting precast elements, they typically involve slotting at the ends of the precast slabs for closely spaced connections, which are most effective in applications where the connection primarily serves to join the ends of slabs. These designs are generally not suitable for wall–slab connections, as they do not provide the necessary load transfer efficiency and structural integrity required at the interface between vertical walls and horizontal slabs.

Building on these studies and aiming to further enhance the connection and construction performance of precast concrete wall panels, a novel slotted concrete composite slab–composite shear wall (SCS-CW) connection is proposed in this study. In this method, slotted concrete composite slabs are used to replace traditional precast slabs with protruding rebars, effectively reducing construction difficulty while ensuring sufficient anchorage length for the reinforcement. The connection presented in this study features lap-spliced reinforcement passing through the post-cast zone, where it works in conjunction with the precast slabs and the composite wall to transfer loads effectively.

The SCS-CW connection is composed of precast slotted concrete composite slabs, precast walls, and post-cast layers. Force transfer between the two precast slotted concrete slabs is achieved through lap-spliced reinforcement placed within the slots. The post-cast layer and the two precast walls together form a bilateral composite shear wall, which connects with the precast slotted concrete slab above, thus creating an integrated connection, as illustrated in Figure 3.

The SCS-CW connection offers significant advantages in terms of constructability and cost-effectiveness. Existing wall–slab connection methods all have varying degrees of limitations in terms of constructability. Taking the composite slabs featuring protruding reinforcement as an example, the factory fabrication necessitates creating openings in the side molds to accommodate the extension of rebars beyond the precast slab edges. Each horizontal reinforcement bar must be threaded through these openings along the entire slab length and subsequently sealed to prevent concrete leakage during casting. This procedure requires stringent precision and substantial manual effort, posing challenges to mechanization. As a result, the production efficiency of precast slabs is low, and manufacturing costs increase. In contrast, the SCS-CW connection requires only simple mold configurations, and on-site construction is simplified to placing lap-spliced reinforcement within the slots followed by casting the post-cast layer. Since the slots are formed during precast fabrication, the number of field rebar alignments and manual corrections is significantly reduced. These improvements can shorten construction cycles, reduce skilled labor demands, and lower the risk of rework due to rebar misalignment.

Table 1. Comparison of different connection methods.

Connection Method	Performance	Constructability	Cost Implications
Proposed slotted system	High ductility and load capacity	Simplified formwork; reduced labor	Cost savings through reduced manual labor
Protruding rebar	Moderate performance	Complex formwork; high labor requirement	Higher due to labor intensity
Mechanical splice systems	High strength and reliability	Requires precise installation	Higher material and labor costs
Grouted sleeve connections	High strength and good ductility	Requires precise installation	Moderate due to material and labor cost

summarizes several state-of-the-art connection methods.

This study involved experimental tests focusing on the static behavior of SCS-CW connections. Four types of connection detailing schemes were designed, and seven full-scale specimens were fabricated, with a focus on examining the reliability of the slotted composite configuration and the influence mechanism of lap-spliced reinforcement within the slots on the mechanical behavior of the connections. Based on the experimental data, calculation methods for the bearing capacity and flexural stiffness of the SCS-CW connections were proposed and further validated through finite element simulations.

2. Overview of the Experimental Program

2.1. Specimen Design

As shown in Figure 4, four types of SCS-CW connection specimens were designed in this study, designated as specimens QB1 to QB4. To minimize experimental randomness and enhance data reliability, two specimens were fabricated for each connection type, except for the cast-in-place control specimen. All specimens shared the same dimensions: 1700 mm in length, 600 mm in width, and 750 mm in height. The precast slab thickness was h₁ = 70 mm, while the post-cast layer measured h₂ = 60 mm. Specimen QB1 served as the control group, representing a fully post-cast slab–composite wall connection. Specimens QB2, QB3, and QB4 were experimental groups with different connection forms. In specimen QB2, no slot was provided, and the longitudinal reinforcement of the precast slab was extended outward. In specimens QB3 and QB4, slots were formed in the precast slab, and lap-spliced reinforcement was placed within the slots, parallel to the longitudinal reinforcement of the precast slab.

According to the recommendations of Jiang et al. [20], for the specimens with slots, the strip slot dimensions were set as follows: depth h₃ = 40 mm, width a = 40 mm, and length n = 190 mm, with a slot spacing m = 150 mm. Three slots were arranged on each side, and the slots were positioned in the horizontal plane to stagger with the longitudinal reinforcement of the precast slab. Lap-spliced reinforcement was placed within the reserved slots at the precast slab, with an anchorage length of 21d, where d refers to the diameter of the lap-spliced reinforcement. In specimen QB3, the lap-spliced reinforcement had the same diameter as the longitudinal reinforcement of the precast slab, whereas in specimen QB4, the lap-spliced reinforcement had a diameter 4 mm greater than that of the longitudinal reinforcement in the precast slab, ensuring that the bonded area of the reinforcement anchored within the slot remained unchanged while reducing the anchorage length to 9.3d. Specimen QB2 had no slots and no lap-spliced reinforcement, with its longitudinal reinforcement directly protruding from the precast slab.

For specimens QB2 through QB4, 8 mm diameter rebars served as the longitudinal reinforcement in the precast slabs, whereas 12 mm diameter rebars were employed in the post-cast layer. The reinforcement spacing in both the precast and post-cast sections was maintained at 150 mm. Transverse reinforcement was provided by ribbed steel bars of the same diameter as the longitudinal reinforcement for confinement. The relevant parameters of each specimen are summarized in

Table 2. Specimen parameters (unit: mm).

Specimen ID	Slab Configuration	Longitudinal Reinforcement in Precast Slab	Lap-Spliced Reinforcement	Slot Dimensions (Depth × Width × Length)
			Layout/Length
QB1	Fully cast-in-place slab	Φ8@150	/	/
QB2	Precast slab with protruding reinforcement	Φ8@150	/	/
QB3	Slotted composite slab	Φ8@150	Φ8@150/12d	40 × 40 × 190
QB4	Slotted composite slab	Φ8@150	Φ12@150/9.3d	40 × 40 × 132

.

2.2. Material Property Tests

The concrete was designed with a strength grade of C40. The specimens were cast in three separate stages. For each stage, three cubic specimens with dimensions of 150 mm were prepared. The standard compressive strength of the concrete was determined by loading the cubes and taking the average value, as summarized in

Table 3. Results of concrete material property tests.

Concrete Type	fc (MPa)			σ/CV	$\overline{{\mathit{f}}_{\mathit{c}}}$ (Mpa)	Ec (Mpa)
	1#	2#	3#
Precast Wall	42.1	43.9	47.5	2.24/5.04%	44.5	35.6
Precast Slab	46.0	48.3	49.1	1.31/2.75%	47.8	32.2
Cast-in-Place Topping Layer	39.3	45.6	44.7	2.78/6.44%	43.2	33.7

Note: σ = standard deviation; CV = coefficient of variation; f_c = measured compressive strength; $\overline{f_c}$ = average compressive strength of concrete cubes; E_c = elastic modulus of concrete.

.

All reinforcement used in the specimens was of HRB400 grade, with diameters of 8 mm and 12 mm. During specimen fabrication, three specimens were cut from the base material of each rebar type for tensile testing. The measured average strength values of each type of reinforcement are summarized in

Table 4. Results of reinforcement material property tests.

D	fy (Mpa)			σ/CV	$\overline{{\mathit{f}}_{\mathit{y}}}$ (Mpa)	fu (Mpa)			σ/CV	$\overline{{\mathit{f}}_{\mathit{u}}}$ (Mpa)	Es (Mpa)
	1#	2#	3#			1#	2#	3#
8 mm	457.8	448.8	447.3	4.64/1.03%	451.3	614.6	616.9	624.3	4.14/0.67%	618.6	2.0E5
12 mm	458.6	462.9	463.3	2.13/0.46%	461.6	631.1	645.1	624.9	8.45/1.33%	633.7	2.0E5

Note: σ = standard deviation; CV = coefficient of variation; D = diameter of reinforcement; f_y = measured yield strength; $\overline{f_y}$ = average yield strength of reinforcement; f_u = measured ultimate tensile strength; $\overline{f_u}$ = average ultimate tensile strength of reinforcement; E_s = elastic modulus of reinforcement.

.

2.3. Loading and Testing Protocol

Since the connection was located in the negative moment region, inverted loading was adopted for convenience, with a support span of 1500 mm. The loading setup is shown in Figure 5. A two-stage control method combining force and displacement control was used in the test: during the force control stage, the load was applied incrementally with 4 kN steps until 80% of the calculated load was reached. Subsequently, the control mode shifted to displacement control, with each increment set equal to the estimated yield displacement, and loading proceeded until the specimen failed. The estimated load values were determined according to the provisions of GB50010-2010 [40]. During the test, two displacement gauges were arranged in the mid-span region to monitor vertical deformation. Furthermore, strain gauges were placed at critical points on both the lap-spliced bars and the reinforcement of the precast slab to quantitatively assess the load transfer behavior between the slotted connection region and the precast slab, as illustrated in Figure 6.

3. Sensitivity Analysis of Influential Parameters

3.1. Experimental Observations

The crack distribution and failure characteristics of the connection specimens for each group are detailed in Figure 7 and Figure 8. All specimens exhibited bending failure modes, with no significant damage observed in the double-sided composite shear walls. Under ultimate conditions, the overall crack distribution in each specimen extended upward from the bottom, with a denser concentration of cracks in the middle region and fewer cracks on the sides. In the later stages of displacement loading, inclined cracks began to develop. In specimens QB2 and QB4, partial separation occurred at the composite interface within the plastic hinge region under the action of bending moments during the later stages of loading. All specimens exhibited significant deformation. After the reinforcement yielded, the displacement–load curves showed a clear plateau in displacement development, followed by a gradual decline in load after reaching the peak, as illustrated in Figure 9. These characteristics indicate that all specimens experienced ductile failure.

When the load reached 16 kN, bending cracks first developed in the plastic hinge region of the post-cast layer in specimen QB1. Cracks appeared in specimens QB2, QB3, and QB4 when the applied load reached 12 kN, 8 kN, and 8 kN, respectively. With increasing load, the number of cracks in the plastic hinge area of the post-cast layer grew progressively, extending vertically upward while their width and length expanded continuously. The longitudinal reinforcement in the post-cast layer within the connection zone yielded at 67.4 kN, with concrete spalling occurring at the edges of the precast slab.

After the load reached its peak bearing capacity, cracks in the plastic hinge region gradually propagated diagonally toward the corners of the connection zone. In specimens QB2, QB3, and QB4, in addition to cracks observed at the splicing locations within the connection zone, multiple cracks also appeared on both sides of the splicing region, indicating that the specimen was subjected to overall flexural behavior. The crack pattern of specimen QB3 closely resembled that of specimen QB1, with no interface separation observed, indicating superior composite performance. The crack development patterns of specimens QB2 and QB4 were similar. Diagonal cracks initiated in the plastic hinge region of the connection at loads of 74.9 kN and 71.6 kN, respectively, and extended horizontally along the composite interface. In the final stages of loading, the diagonal crack merged with a neighboring vertical crack and continued to advance toward the corner area of the connection.

A comparison between specimens QB2 and QB3 revealed that at failure, the crack width at the splicing location in QB3 was significantly smaller than that in QB2, and no cracks were observed at the composite interface. This indicates that the use of lap-spliced reinforcement within the slots of the composite slab effectively controlled crack development in the plastic hinge region and provided better composite performance. Comparing specimens QB3 and QB4 further demonstrated that the anchorage length of the lap-spliced reinforcement had a significant influence on the connection effectiveness between slotted slab elements.

3.2. Load–Displacement Curves

Figure 9 shows the load–displacement relationships for all specimens. Specimen QB1 exhibited the best deformation performance; in the later stage of loading, the load showed no significant decline with increasing displacement. The peak load-carrying capacities of the four groups of specimens were relatively close. Specimen QB1 had the highest peak load of 108 kN, while the peak loads of specimens QB2, QB3, and QB4 were 98 kN, 99 kN, and 100 kN, respectively. Specimens QB1, QB2, and QB3 all exhibited good ductility after yielding, with specimen QB3 achieving a higher ultimate displacement than specimen QB2. The ductility of specimen QB4 was relatively poor, primarily because the anchorage length of the lap-spliced reinforcement within the slots was shorter, resulting in a weaker connection between the slotted slab elements and insufficient force transfer.

Based on the experimental observations and the load–displacement curves, it is evident that the SCS-CW connections experienced a three-stage response from loading to failure, consistent with that of cast-in-place connections. The first stage was the elastic behavior stage, during which the structural behavior conformed to elastic theory, no visible cracks appeared, and the reinforcement and concrete in the connection zone between the precast slab and the post-cast layer worked together to ensure effective force transfer. The second stage was the cracked working stage, during which cracks began to develop and propagate in the connection zone, resulting in decreased stiffness, while the structure continued to retain a degree of load bearing and deformability. The third stage began with the yielding of the reinforcement and continued until the component reached its ultimate state and eventually collapsed.

Table 5. Key mechanical performance indicators of each specimen.

Specimen ID	Mcr (kNm)	My (kNm)	θcr (%)	dy (mm)	Mu (kN)	df (mm)	μ	Bcr (kNm2)	By (kNm2)
QB1	5.2	21.9	0.103	3.8	108.4	/	/	3060.5	1229.9
QB2	3.9	24.3	0.086	6.3	97.1	74.6	11.8	2560.7	1146.1
QB3	2.6	24.9	0.058	6.8	98.6	83.1	12.2	2520.1	1124.4
QB4	2.6	23.3	0.057	4.5	99.7	28.5	6.3	2543.6	1139.3

summarizes the key mechanical performance indicators of each specimen during loading, including the cracking load M_cr, yield load M_y, yield displacement d_y, peak load M_u, failure displacement d_f, ductility coefficient μ, cracking stiffness B_cr, and yield stiffness B_y. The ductility coefficient is calculated as the ratio of the d_f to d_y. The cracking stiffness refers to the initial stiffness exhibited prior to crack initiation, while the yield stiffness represents the stiffness observed between the onset of cracking and the yielding of the reinforcement. The formula for calculating the cracking stiffness is shown in Equation (1).

$$B_{cr}={\displaystyle \frac{M_{cr}}{{\theta }_{cr}}}$$

(1)

As shown in the table, specimen QB1 had a relatively small cracking load P_cr, but a large cracking stiffness B_cr. The cracking stiffness of specimens QB3 and QB4 was slightly lower than that of specimen QB2, indicating that introducing rectangular slots in the precast slab can reduce the cracking stiffness of the connections. The ratios of the ultimate load-carrying capacities of specimens QB2, QB3, and QB4 to that of specimen QB1 were 0.91, 0.92, and 0.93, respectively, suggesting that the structural capacity of the precast connections approached that of specimen QB1, and that the use of lap-spliced reinforcement at the slotted slab ends effectively enabled force transfer between components. Specimen QB3 exhibited a greater ductility coefficient compared to QB2, suggesting that compared to using protruding reinforcement, setting lap-spliced reinforcement within slots can improve the ductility of the connection. The ductility coefficient of specimen QB3 was also much higher than that of specimen QB4, demonstrating that the anchorage length of the lap-spliced reinforcement is a critical parameter affecting connection ductility. Moreover, although the ductility coefficient of specimen QB4 was lower than that of QB2 and QB3, it still reached 6.3 and exhibited a clear yielding stage, indicating that SCS-CW connections possess strong deformation capacity under different lap reinforcement anchorage lengths and can meet the requirements for engineering applications.

3.3. Reinforcement Strain Results

The strain–displacement curves were obtained at key measurement points for each specimen, as depicted in Figure 10. The strain of both the lap-spliced reinforcement and the mid-span reinforcement in the post-cast layer was obtained by averaging readings from six measurement points each. The horizontal axis represents the mid-span displacement of the specimen under loading. The horizontal dashed line in the figure represents the yield strain of the reinforcement, calculated by dividing f_y by E_s, resulting in a yield strain of ε_y = 2300 με.

During the elastic stage of each specimen, the mid-span reinforcement in the cast-in-place topping layer was subjected to tension, and its strain increased rapidly with the increase in mid-span displacement. In the early loading stage of specimen QB2, the protruding reinforcement from the precast slab was subjected to compression, and its strain increased continuously with the increase in mid-span displacement. As the mid-span reinforcement in the post-cast layer approached yielding, the protruding reinforcement in the precast slab gradually shifted from a compressive state to a tensile state, with its strain further increasing as the load increased. For specimens QB3 and QB4, the lap-spliced reinforcement within the slab slots was under tension throughout the entire loading process. After the mid-span reinforcement in the post-cast layer yielded, the strain growth rate of the lap-spliced reinforcement in the precast slab slots significantly accelerated with increasing mid-span displacement until the specimen failed. At failure, the protruding reinforcement in specimen QB2 and the lap-spliced reinforcement in specimen QB3 both reached yielding, while the lap-spliced reinforcement in specimen QB4 had not yet yielded.

The experimental results showed that both the protruding reinforcement from the precast slab and the lap-spliced reinforcement within the slab slots began to carry the main tensile force after the mid-span reinforcement in the cast-in-place topping layer yielded in tension. This suggests that both the protruding and lap-spliced reinforcements in the precast slab contributed significantly to improving the component’s ductility, while only slightly affecting the peak strength in the early phase of loading. A comparison between specimens QB3 and QB4 reveals that, once the mid-span reinforcement in the cast-in-place topping layer yielded, the strain in the lap-spliced reinforcement of both specimens escalated rapidly with increasing displacement. However, in specimen QB4, the component failed before the lap-spliced reinforcement reached yielding, and its mid-span displacement was lower than that of specimen QB3. This suggests that due to the shorter anchorage length, the lap-spliced reinforcement in specimen QB4 did not fully contribute before the longitudinal reinforcement in the post-cast layer entered the strength degradation stage, resulting in lower ductility for specimen QB4 compared to QB3. Therefore, for SCS-CW connections, ensuring sufficient anchorage length within the slot is essential to fully mobilizing the lap-spliced reinforcement and achieving reliable force transfer connections.

4. Discussion of the Design Method

Based on the experimental results, all connection specimens exhibited a favorable load-bearing performance, and the overall failure behavior corresponded to that of flexural failure in beam-like members. Therefore, the fundamental design principles for RC bending members may serve as a reference for developing the design approach for SCS-CW connections.

4.1. Flexural Load-Carrying Capacity Calculation

Specimen QB1 is a fully cast-in-place slab and can therefore be directly analyzed using the formulas for the bending capacity of doubly reinforced rectangular sections provided in GB50010-2010 [40], as shown in Equations (2) and (3).

$${\alpha }_1{f}_{\mathrm{c}}bx+f_{\mathrm{y}}^{\prime }{A}_{\mathrm{s}}^{\prime }=f_y{A}_s$$

(2)

$$M_{\mathrm{u}}={\alpha }_1{f}_{\mathrm{c}}bx(h_0-{\displaystyle \frac{x}{2}})+f_y^{\prime }{A}_s^{\prime }(h_0-a_s^{\prime })$$

(3)

where ${\alpha }_1$ is related to the concrete strength and is typically taken as 1.0, $b$ is the width of the cross-section, $x$ is the depth of the compression zone, $f_y^{\prime }$ denotes the yield strength of the compressive reinforcement, and $A_s^{\prime }$ is its area. $A_s$ represents the area of the tensile reinforcement, $h_0$ is the effective depth of the section, $a_s^{\prime }$ is the distance from the compressive reinforcement to the compressive edge, and $M_u$ is the flexural capacity of the cross-section.

The experimental results indicate that, at the time of failure of specimen A, the tensile reinforcement had already entered the strain-hardening stage. Therefore, when using Equation (1) to calculate the depth of the compression zone x, it is recommended to replace f_y in the code formula with the f_u of the reinforcement to more accurately reflect the actual stress state of the member.

By substituting the measured f_c, f_u, and f_y^’ into the formula, the theoretical bending capacity of the specimen is calculated as M_u = 33.61 kN·m, corresponding to a load of 103.4 kN.

Specimens QB2, QB3, and QB4 exhibited typical properly reinforced failure modes during the tests, characterized by the yielding of the tensile reinforcement in the cast-in-place topping layer and the crushing of the concrete in the compression zone. No separation at the composite interface occurred before reaching the peak load. Therefore, for these three types of precast specimens, the flexural capacity can still be calculated using the rectangular stress block method recommended in GB50010-2010 [40].

Moreover, according to JGJ1-2014 [41], when calculating the performance of connections in prefabricated monolithic structures, the calculation can be based on the methods for the normal sections of concrete members, considering only the reinforcement that crosses the section and is reliably anchored.

Accordingly, when calculating the flexural capacity of the connection sections for the three groups of specimens, the contribution of the compressive reinforcement in the precast slab can be conservatively neglected. Thus, the bending capacity can be calculated using Equations (4) and (5).

$${\alpha }_1{f}_{\mathrm{c}}bx=f_{\mathrm{y}}{A}_{\mathrm{s}}$$

(4)

$$M_{\mathrm{u}}={\alpha }_1{f}_{\mathrm{c}}bx(h_0-{\displaystyle \frac{x}{2}})$$

(5)

The bending capacities of the precast specimens were calculated using Equations (4) and (5) and compared with the test results, as shown in

Table 6. Comparison of load-carrying capacity (unit: kN).

Specimen ID	Pe (kN)	Pt (kN)	Pe/Pt
QB1	108.4	103.4	0.96
QB2	97.1	96.3	0.99
QB3	98.6	96.3	0.98
QB4	99.7	96.3	0.97

Note: P_e = experimental value; P_t = theoretical value.

. The ratios of the theoretical values to the experimental values for specimens QB1 to QB4 were 0.96, 0.99, 0.98, and 0.97, respectively, with errors not exceeding 10%. This indicates that the load-carrying capacity of slotted concrete composite slab–wall connections can be accurately calculated using the normal section design method for concrete members. The calculated results are generally consistent with the experimental measurements and are conservatively safe, making them applicable for engineering design.

4.2. Flexural Stiffness Calculation

GB50010-2010 [40] provides the formulas for the flexural stiffness before cracking B₀ and the short-term stiffness after cracking B_s for RC bending members, as shown in Equations (5) and (6). DG/TJ08-2071-2010 [42] provides the short-term stiffness calculation method for precast concrete members, as illustrated in Equation (7).

$$B_0=0.85E_{\mathrm{c}}{I}_0$$

(6)

$$B_{\mathrm{s}1}={\displaystyle \frac{E_{\mathrm{s}}{A}_{\mathrm{s}}{h}_0^2}{1.15\psi +0.2+\frac{6\alpha {\rho }_{\mathrm{E}}\rho }{1+3.5{\gamma }_f^{\prime }}}}$$

(7)

$$B_{\mathrm{s}2}={\displaystyle \frac{E_{\mathrm{s}}{A}_{\mathrm{s}}{h}_0^2}{0.7+0.6\frac{h_1}{h_2}+\frac{4.5{\alpha }_{\mathrm{E}}\rho }{1+3.5{\gamma }_{\mathrm{f}}^{\prime }}}}$$

(8)

where $E_s$ is the elastic modulus of the reinforcement, $E_c$ is the elastic modulus of the concrete, ${\alpha }_E$ is the ratio of the elastic modulus of the reinforcement to that of the concrete, $h_2$ is the height of the specimen, $h_1$ is the height of the precast slab, $\psi$ is the longitudinal tensile reinforcement strain non-uniformity factor between cracks, $I_0$ is the transformed section moment of inertia, and $\rho$ is the longitudinal tensile reinforcement ratio. Since the specimens have no tensile flanges or webs, the term ${\gamma }_f^{\prime }$ is not considered in the calculation.

By substituting the experimental data, the flexural stiffness of the components can be calculated and compared with the test results, as illustrated in

Table 7. Comparison of flexural stiffness (unit: kN·m²).

Specimen ID	${\mathit{B}}_{\mathit{c}\mathit{y}}$(kNm2)	${\mathit{B}}_0$(kNm2)	${\mathit{B}}_{\mathit{c}\mathit{y}}$$/{\mathit{B}}_0$	${\mathit{B}}_{\mathit{y}}$(kNm2)	${\mathit{B}}_{\mathit{s}1}$(kNm2)	${\mathit{B}}_{\mathit{y}}$$/{\mathit{B}}_{\mathit{s}1}$	${\mathit{B}}_{\mathit{s}2}$(kNm2)	${\mathit{B}}_{\mathit{y}}$$/{\mathit{B}}_{\mathit{s}2}$
QB1	3060.5	3277.5	0.93	1229.9	1021.5	1.20	986.0	1.24
QB2	2560.7		0.78	1146.1		1.12		1.16
QB3	2520.1		0.77	1124.4		1.10		1.14
QB4	2543.6		0.78	1139.3		1.12		1.16

. The results show that for specimen QB1, the cracking stiffness calculated using Equation (5) differed from the measured stiffness by about 7%, whereas for specimens QB2, QB3, and QB4, the calculation error exceeded 20%, indicating that initial defects existed in the precast connections, and they cannot be designed fully according to the behavior of uncracked RC members. However, the short-term stiffness derived from Equations (6) and (7) showed strong consistency with the tests, with the calculated-to-measured stiffness ratios for specimens QB2, QB3, and QB4 ranging from 1.10 to 1.16.

Overall, for the SCS-CW connections studied in this paper, the short-term stiffness of the components can be calculated using the formulas provided in GB50010-2010 [40] and the Code for DG/TJ08-2071-2010 [42], and the calculation results are conservatively safe.

5. Finite Element Model (FEM)

5.1. Model Parameters

To further investigate the key parameters affecting SCS-CW connections, a detailed numerical model was established in ABAQUS, as shown in Figure 11. The reinforcement, precast concrete, and cast-in-place topping concrete were modeled independently. Reinforcement was modeled using T3D2 elements, while concrete was modeled using C3D8R elements. The interaction between the rebars and concrete was assumed to be perfect, and the rebars were embedded into the concrete. A uniform mesh size of 20 mm was assigned to both the concrete and the reinforcement. Contact interactions were established between the concrete and the supports. For normal contact behavior, “hard contact” was applied to ensure stress transmission; for tangential contact behavior, frictional contact was used with a friction coefficient of 0.6 [43].

Interfaces between the precast and post-cast concrete were present in the contact region, significantly influencing the bearing capacity and failure modes of the connections. To simulate the interfacial behavior between the new and existing concrete, a cohesive model was adopted, with the corresponding parameters provided in

Table 8. Cohesive model parameters.

μ	Pressure–Overclosure	Stiffness (MPa/mm)			Peak Stress (mm)
		Knn	KSS	Ktt	${\mathit{t}}_{\mathit{n}}^0$	${\mathit{t}}_{\mathit{s}}^0$	${\mathit{t}}_{\mathit{t}}^0$
0.6	“Hard” Contact	1 × 105	10	10	1.6	0.8	0.8

. In ABAQUS, six degrees of freedom were defined for the boundary conditions: UR1, UR2, and UR3 for rotational degrees of freedom; U1, U2, and U3 for translational degrees of freedom. Since the experimental boundary condition was simply supported, the boundary conditions for the two supports were set as UR1 = UR2 = UR3 = 0, and U1 = UR1 = UR2 = UR3 = 0, and the horizontal restraint at the loading point at the base of the shear wall was released. The vertical load in the finite element analysis was applied at a coupling point at the end of the wall, consistent with the loading condition in the experiment.

5.2. Material Properties

In this study, the concrete material was modeled using the Concrete Damaged Plasticity (CDP) model [44,45], with its stress–strain behavior defined based on the equations specified in the Chinese design code GB50010-2010 [40], with a Poisson’s ratio of 0.2. The initial tangent elastic modulus and plasticity parameters were determined according to reference [46], and the damage factors were set based on reference [47], as shown in

Table 9. Constitutive parameters of concrete.

Eccentricity	Dilation Angle	fbo/fco	K	Viscosity Parameter	$\mathsf{\nu }$
0.1	30	1.16	0.667	0.0005	0.3

.

Reinforcement was modeled using an elastic–plastic hardening model, with the von Mises yield criterion. The elastic modulus E_s was taken as 2 × 10⁵ N/mm, and Poisson’s ratio was taken as 0.3. The stress–strain curve was bilinear, featuring a mild slope after yielding, with the post-yield stiffness taken as 0.01 E_s.

5.3. Finite Element Validation

In this study, specimens QB1 and QB3 were selected for model validation. The overall failure modes of specimens QB1 and QB3 obtained from the FEM analysis are illustrated in Figure 12. The two specimens exhibited similar failure modes. The composite shear wall and the support regions at the ends of the slab remained in the elastic stage. Concrete tensile damage mainly occurred near the mid-span connection zone, where plastic hinges formed on both sides of the connection. Compressive damage occurred at the ends of the precast slab near the connection zone.

The failure modes of the reinforcement and the slots are shown in Figure 13. Both the reinforcement in the post-cast layer and the lap-spliced reinforcement at the slot locations reached yielding, indicating that the lap-spliced reinforcement was fully mobilized. No interface separation was observed, suggesting good composite action and force transfer.

The failure modes derived from the FEM analysis closely aligned with the experimental observations. As shown in Figure 14, the load–displacement curve obtained from the FEM analysis agrees well with the experimental results in terms of stiffness, bearing capacity, and ductility.

5.4. Parametric Analysis and Design Recommendations

According to the experimental results, in specimen QB4, although the bonded area between the reinforcement and concrete within the slots was kept constant by increasing the diameter of the lap-spliced reinforcement and shortening its anchorage length, the ductility was significantly reduced, and the lap-spliced reinforcement remained unyielding in subsequent loading phases. This suggests that the configuration of the lap-spliced reinforcement within the slots is a key parameter affecting the performance of the connection specimens. To explore the optimal lap-spliced reinforcement configuration, this study quantitatively analyzed the effects of the reinforcement quantity, diameter, and anchorage length on the mechanical performance of the connections in ABAQUS.

To change the quantity of lap-spliced reinforcement, the spacing of the slots must be adjusted. The displacement–load curves of the finite element models under different slot spacings are illustrated in Figure 15a. Reducing the slot spacing (i.e., increasing the quantity of lap-spliced reinforcements) can enhance the ductility of the component; however, when the slot spacing is reduced below 120 mm, the improvement in ductility becomes less significant. In comparison to the specimen featuring a 200 mm slot spacing, the initial stiffness decreased by 6.4%, 17.1%, and 22.7% for slot spacings from 150 to 100 mm, while the peak load-carrying capacity increased by 1.7%, 6.5%, and 7.2%. The reason is that increasing the number of slots introduces more initial defects at the connection, weakening the initial stiffness, whereas increasing the quantity of lap-spliced reinforcement enhances the reinforcement ratio in the tensile zone, thereby improving the bearing capacity of the connection. Therefore, a slot spacing of 120 mm is recommended.

The displacement–load curves from finite element models featuring varying diameters of lap-spliced reinforcement are presented in Figure 15b. Increasing the diameter of the reinforcement enhances the component’s ductility. However, when the reinforcement diameter exceeds 10 mm, the contribution of the reinforcement to the ductility becomes less significant because the reinforcement in the cast-in-place topping layer has already reached its ultimate strength. The increase in the diameter of the lap-spliced reinforcement enhances the amount of reinforcement in the tension zone, and compared to 6 mm, the bearing capacities for reinforcement with diameters ranging from 8 to 12 mm increased by 4.7%, 9.8%, and 12.8%, respectively. It is recommended that the slot reinforcement adopt a diameter greater than 10 mm to fully ensure the connection’s ductility.

The displacement–load curves of the finite element models for lap-spliced reinforcements with different anchorage lengths are presented in Figure 15c. (The slot length is the same as the anchorage length of the lap-spliced reinforcement.) When the anchorage length of the lap-spliced reinforcement was less than 21d, the ductility of the connection significantly decreased. This is because shortening the anchorage length reduces the bond strength of the lap-spliced reinforcement, leading to slippage and an uneven force distribution. However, an excessively long slot length also increases the initial defects in the connection. Compared to an anchorage length of 14d, the initial stiffness decreased by 5.1%, 6.2%, and 6.8% for anchorage lengths of 17d, 21d, and 28d, respectively. Therefore, an anchorage length of 21d is recommended, as it ensures sufficient bond strength of the reinforcement without excessively increasing the initial defects in the connection.

Based on the parametric analysis conducted in ABAQUS, the optimal values for key design parameters are summarized in

Table 10. Design recommendations.

Key Parameter	Slot Spacing	Diameter of the Lap-Spliced Reinforcement	Anchorage Length of the Lap-Spliced Reinforcement
Recommended Value	120 mm	10 mm	21d

. These recommendations are intended to ensure that the SCS-CW connection achieves sufficient ductility without significantly compromising its initial stiffness. The proposed values are applicable to precast slabs with a thickness of no less than 70 mm, which ensures adequate concrete cover beneath the slot to allow the lap-spliced reinforcement to fully develop bond strength and transfer forces effectively. It should be noted that the recommended anchorage length assumes the use of straight lap-spliced bars. If hooked bars are used instead, the anchorage behavior will differ, and the suggested values may no longer be applicable. This condition warrants further investigation and can be addressed in future studies.

6. Discussion

Although the SCS-CW connection demonstrated favorable mechanical performance under static loading conditions, and the finite element simulation results showed good agreement with the experimental data, several limitations remain in the present study:

1. The current study focused exclusively on the static performance of slotted concrete composite slab–composite wall connections under monotonic loading. While the obtained results provide valuable insights into the mechanical performance of the proposed connection, seismic and dynamic demands are of critical importance for practical applications, especially in regions with high seismic risk. Wall–slab connections are known to experience cyclic loading and repeated stress reversals during earthquakes, which can significantly affect their stiffness degradation, energy dissipation capacity, and cumulative damage behavior. Therefore, further experimental investigations under cyclic or hysteretic loading are essential to comprehensively evaluate the seismic performance and structural resilience of the proposed system.

2. In ABAQUS, a key assumption is that the interaction between reinforcement and concrete is modeled using embedded region constraints, which idealize the bond as perfect and neglect the relative slip at the interface. The absence of an explicit bond–slip model may lead to a slight overestimation of stiffness and an underestimation of local deformation or slippage at the reinforcement–concrete interface. In practice, these effects could influence crack development, load transfer efficiency, and ductility.

3. This research was primarily concerned with the short-term mechanical response of the proposed slotted connection under static loading. Time-dependent effects such as creep, shrinkage, and temperature-induced deformations were not considered. These long-term behaviors may alter the internal stress distribution, deformation compatibility, and serviceability of precast composite systems. Future research will incorporate long-term performance evaluations through extended numerical modeling and experimental studies to assess the time-dependent behavior of the connection under realistic service conditions.

4. Although precast concrete connections offer advantages such as consistent quality and environmental friendliness during construction, it should be noted that the present study still utilizes conventional concrete materials, which do not fully align with the principles of sustainable development. Future research should aim to incorporate low-carbon cementitious materials [48,49] and carbonation-based maintenance strategies [50,51] to enhance the environmental performance of the proposed system and move toward more sustainable precast construction practices.

7. Conclusions

Experimental research was conducted into the static performance of SCS-CW connections in this study. The key findings are summarized as follows:

(1) All specimens exhibited bending failure during the experiments. No separation at the composite interface was observed at the final failure stage of specimen QB3, and its failure mode was generally consistent with that of specimen QB1. This indicates that the connection constructed with slotted composite slabs achieved good composite action.

(2) The ductility coefficient of specimen QB3 was higher than that of specimen QB2, demonstrating that slotted concrete composite slab–wall connections possess favorable ductility. The ductility coefficient of specimen QB4 was relatively lower, mainly because the anchorage length of the lap-spliced reinforcement within the slots was shortened, resulting in a weaker connection between the slotted slab elements and insufficient force transfer.

(3) Under ultimate conditions, the contribution of compressive reinforcement in the precast slab can be neglected when calculating the flexural capacity of SCS-CW connections. The calculated values closely match the experimental results and tend to be conservative, making them suitable for engineering design.

(4) The short-term stiffness of SCS-CW connections was generally consistent with that of connections using protruding reinforcement, and it can be calculated using the formulas provided in the code, with conservatively safe results.

(5) Parametric finite element analysis showed that increasing both the number and diameter of lap-spliced reinforcement improves the connection’s ductility and load capacity; however, too many slots can reduce the connection’s stiffness. When the anchorage length of the lap-spliced reinforcement was less than 21d, the ductility of the connection decreased significantly. The anchorage length of 21d was optimal.

(6) The slotted concrete composite slab–composite shear wall connection features a simple configuration and high assembly efficiency while ensuring good mechanical performance. It can also effectively reduce construction costs, improve construction efficiency, and decrease on-site labor, offering broad application prospects.