Experimental Investigation on the Seismic Performance of Novel Prefabricated Composite RC Shear Walls with Concrete-Filled Steel Tube Frame

Abstract

The present study proposed novel prefabricated composite RC shear walls with a concrete-filled steel tube frame (CCRCSW-CFST) because of the superior seismic performance of shear walls incorporating CFSTs as boundary-restrained members. One cast-in-place reinforced concrete shear wall (RCSW) and seven CRCSW-CFSTs, each varying in axial compression ratios, concrete strengths, and shear span ratios, were designed for experimental analysis. Cyclic loading tests were performed on these specimens, yielding the following results: (1) Compared to reinforced concrete shear walls, CCRCSW-CFSTs demonstrated superior seismic performance, with 14.2% increased ductility and 47.5% greater energy dissipation capacity. (2) Elevating the axial compression ratio in CCRCSW-CFSTs resulted in increased yield strength, peak strength, and stiffness. Conversely, this adjustment also expedited the degradation of stiffness with displacement and decreased both ductility and ultimate deformation. (3) The peak displacement and ultimate displacement of CCRCSW-CFSTs were both increased with an increase in concrete strength. Increasing the axial compression ratio enhanced the initial stiffness of CCRCSW-CFSTs and mitigated the rate at which stiffness deteriorated with increasing displacement. (4) The stiffness, peak and ultimate displacements, peak and ultimate loads, and shear span ratio of CCRCSW-CFSTs were significantly reduced as the shear span ratio was increased. (5) The minor slip between the reinforced concrete panel of the precast slab and the encasing C-shaped steel contributed to an increase in early-stage energy dissipation of the CCRCSW-CFSTs.

1. Introduction

Reinforced concrete shear walls play a crucial role in the resistance of lateral forces in high-rise buildings. However, conventional RC shear walls may be insufficient in providing the necessary seismic strength, especially on lower floors of high-rise and multi-story structures, where axial pressures and bending moments are significantly higher [1,2]. An increase in wall thickness and densified reinforcements are the most prevalent approaches to satisfying seismic design requirements [3,4]. However, these procedures have the potential to complicate construction, increase the self-weights of buildings, and reduce the usable building areas [5].

Researchers have developed numerous composite shear wall systems by incorporating structural steel elements to improve seismic performance, thereby overcoming the constraints of RC shear walls. Previous research has identified the following primary categories of composite shear walls: single-sided or double-sided steel-plate composite shear walls [6,7], embedded steel-truss composite shear walls [8,9], embedded steel-frame composite shear walls [10,11,12], encased steel–concrete composite shear walls [13,14], concrete-filled steel tube composite shear walls [15], and composite shear walls with embedded X-bracing [16]. These studies have established that the incorporation of structural steel into reinforced concrete shear walls can substantially enhance their load-bearing capacity and ductility, resulting in satisfactory results [17,18].

The shear wall boundary element is a critical factor that affects seismic performance. Incorporating structural steel into this zone significantly enhanced the composite shear wall’s seismic performance [19]. Recent research has demonstrated that shear walls with structural steel in their boundary element can endure higher overturning moments and initial stiffness than conventional RC shear walls [20,21]. Furthermore, these steel-reinforced shear walls offer enhanced lateral load-bearing capacity and improved energy dissipation [22,23]. The concrete in the limbs of reinforced concrete shear walls has limited tensile strength and is prone to crushing under vertical and horizontal loads [24], a problem that cannot be fully resolved by simply incorporating structural steel. Concrete-filled steel tubes (CFSTs) offer a solution to the issue of limited tensile strength in reinforced concrete shear walls by serving as boundary elements. Their confining effect on concrete not only addresses the problem but also significantly enhances the seismic performance of shear walls [25]. The cooperation between CFST columns and concrete webs has been demonstrated in numerous studies [26,27,28], resulting in the effective management of structural deformations and the enhancement of energy dissipation capacity [29].

Furthermore, it has been demonstrated that shear walls with CFST boundary elements offer remarkable seismic performance [30]. However, connecting CFST columns to reinforced concrete webs presents new challenges, requiring a complex and labor-intensive process of welding shear keys onto steel tubes and integrating them with the wall panels during casting. Therefore, this research introduces a prefabricated composite shear wall with a CFST frame to overcome these limitations. As illustrated in Figure 1, the design integrates a CFST frame with prefabricated web panels. The prefabricated web panel was composed of a reinforced concrete slab encased in C-shaped steel sections with a top box beam. These steel sections on the sides were welded to square steel tubes to form vertical seams, while those at the bottom of the upper web panel were welded to the top box beams of the lower web panels, forming horizontal seams. Fill plates were mounted on the exterior of the prefabricated web panels to ensure precise spacing between the panels and the steel tubes, which was essential for achieving the correct weld thickness and quality. The lower flanges of the box-shaped beams at the top of the prefabricated web panels were designed to extend beyond the wall surfaces, offering temporary support for the prefabricated floor slabs. Steel tubes were assembled in batches of two to three stories, with concrete poured into the tubes following the installation of the prefabricated web panels. The benefits of this system are as follows: (1) CFST frames enhance seismic performance and reduce the risk of collapse; (2) welded connections require less precision and simpler assembly, thereby improving efficiency; and (3) box beams can be adapted to connect floor slabs, which increases assembly rates.

2. Experimental Designs

2.1. Specimen Design

Seven new CRCSW-CFST specimens and one RCSW specimen were designed and fabricated following current standards [31,32]. Figure 2 shows the construction details of the specimens, while

Table 1. Specimen parameters.

Specimens	Shear Span Ratio	Axial Compression Ratio	Type of Concrete
RC-2.0-0.3-C50	2.0	0.3	C50
CFST-2.0-0.3-C50	2.0	0.3	C50
CFST-2.0-0.2-C50	2.0	0.2	C50
CFST-2.0-0.4-C50	2.0	0.4	C50
CFST-2.0-0.3-C40	2.0	0.3	C40
CFST-2.0-0.3-C60	2.0	0.3	C60
CFST-1.5-0.3-C50	1.5	0.3	C50
CFST-2.5-0.3-C50	2.5	0.3	C50

The naming scheme of the specimens is as follows: the first component denotes the specimen type: “RC” for conventional RCSW specimens and “CFST” for novel CRCSW-CFST specimens; the second component represents the shear span ratio; the third component represents the axial compression ratio; the fourth component specifies the concrete strength of the specimen.

summarizes the detailed parameters. The specimens were developed to compare the advantages of CRCSW-CFSTs with traditional RCSWs and to evaluate the influence of axial compression ratio, concrete strength, and shear span ratio on the performance of the developed CRCSW-CFSTs. Vertical and horizontal rebars with a diameter of Ø6/100 mm were used to reinforce the concrete webs in all specimens. The steel content in the CRCSW-CFST boundary elements was comparable to that in the RCSW boundary elements. The reinforcement ratios for the RCSW boundary elements were determined using the equivalent strength principle, aligning with the steel quantities used in the CRCSW-CFST boundary elements. Samples with varying shear span ratios showed varying heights. For instance, the CFST-2.0-0.3-C50, CFST-1.5-0.3-C50, and CFST-2.5-0.3-C50 specimens demonstrated heights of around 1550, 1125, and 1975 mm, respectively, while the height of other specimens was determined to be 1550 mm.

The RCSW specimens were cast monolithically, while the rigid reinforced concrete foundation (including steel–concrete columns), shear walls, and CRCSW-CFST specimen beams were prefabricated separately. The installation sequence proceeded as follows: First, the prefabricated web panels were placed between the steel–concrete columns of the rigid foundation. Next, the bottom and sides of the shear walls were welded, finalizing the installation of the web panels. Reinforcement bars were then secured in place, and concrete was poured over the upper sections of the prefabricated web panels and the CFST frame, thus completing the loading beam construction.

2.2. Materials Properties

(1)

Concrete

Concrete strength grades for the specimens were determined to be C40, C50, and C60, corresponding to characteristic cube compressive strengths of 40, 50, and 60 MPa, respectively, following existing standards [33]. Three cubic specimens with dimensions of 150 mm × 150 mm × 150 mm were assessed for each grade.

Table 2. Results obtained from concrete coupon test.

Types of Concrete	Ultimate Pressure $\mathit{P}$ (kN)	Compressive Strength ${\overline{\mathit{f}}}_{\mathit{c}\mathit{u},\mathit{k}}$ (MPa)	Axial Compressive Strength ${\mathit{f}}_{\mathit{c}\mathit{k}}$ (MPa)
C40	981	43.6	29.16
C50	1329	59.10	38.24
C60	1533	68.12	43.72

summarizes the average results of each strength grade [34,35].

(2)

Steels

The specimens were reinforced with HRB400-type reinforcing bars, which included both longitudinal bars and stirrups. The steel plates for the C-shaped sections and box beams in the prefabricated web panels, as well as the steel tubes used in the CFST columns, were of the Q235B type. Steel bars and plates of varying diameters and thicknesses were subjected to monotonic tensile tests, with three specimens per specification [36,37]. The average test results are summarized in

Table 3. Coupon test results.

Types	Yield Strength (MPa)	Ultimate Strength (MPa)	Elastic Modulus (×105 MPa)	Yield Strain (×10−6)
Bar Ø 6	532	695	1.99	2673
Bar Ø 14	395	576	2.03	1945
2 mm	272	384	2.06	1175
5 mm	330	375	2.03	1856

.

2.3. Loading Device

The experimental setup for applying axial and cyclic lateral loads is illustrated in Figure 3. A 300-ton jack attached to the reaction frame through a sliding support to accommodate free movement during loading applied the vertical load. A highly rigid distribution beam was placed between the lower end of the jack and the loading beam to ensure uniform application of the vertical load across the specimens. Horizontal load was applied using a 600 kN MTS electro-hydraulic servo actuator, which was fixed to the reaction wall on one end and connected to the loading beam via a high-strength lead screw on the other. Unidirectional hinges at both ends of each actuator ensured the stable application of horizontal loads. The foundation beams of the specimens were anchored to the ground using a pressure beam and anchor bolts, with a horizontal jack placed between the foundation beam and the limit blocks to prevent sliding.

2.4. Loading Protocol

(1)

Vertical load

A vertical load was first applied to the specimens using a vertical jack, with the magnitude of the applied load ($N_v$) determined using Equation (1).

$$N_v=f_{ck}·A·n$$

(1)

where $n$ represents the axial compression ratio; $f_{ck}$ shows the concrete axial compressive strength; and $A$ denotes the shear wall specimen cross-sectional area.

(2)

Horizontal load

After applying the vertical load, a horizontal load was introduced to the specimen. According to the standards [38], the horizontal load was applied in two stages: force control and displacement control, as shown in Figure 4 [39]. The force control stage commenced with an initial load of 0 kN, which was gradually increased by 5 kN increments. Each increment was applied once and cycled at a rate of 5 kN/s until either the RCSW demonstrated cracking or the steel pipe in the CRCSW-CFST yielded. In the displacement control stage, the amplitude was adjusted to the maximum displacement observed in the previous stage. Each displacement level was cycled thrice at a rate of 1 mm/s until the specimen load capacity dropped below 85% of the peak load, at which point the specimen was considered to have reached failure [40].

2.5. Measuring Device

The measurement setup comprised dial gauges (D1–D4) and wire displacement transducers (CP1–CP4), as illustrated in Figure 5. D1–D3 were used to detect lateral movement or uplift of the rigid foundation, while D4 measured the horizontal displacement of the loading beam. CP1 and CP4, placed on the exterior sides of the shear wall, recorded vertical deformations, and CP2 and CP3, positioned diagonally across the wall, measured diagonal deformations. The bending deformation was calculated using Equation (2), while the shear deformation was determined by Equation (3) [41].

$${\gamma }_c=h\times \Phi ={\displaystyle \frac{{\delta }_3-{\delta }_2}{l}}$$

(2)

$${\gamma }_s={\displaystyle \frac{{∆}_s}{h}}={\displaystyle \frac{{∆}_{s1}+{∆}_{s2}}{2h}}={\displaystyle \frac{d\left({\delta }_4-{\delta }_1\right)}{2hl}}$$

(3)

where ${\gamma }_c$ and ${\gamma }_s$ denote the bending and shear deformation angles, respectively; ${∆}_s$ represents the shear deformation displacement, indicated by D4; $\Phi$ shows the average curvature of the wall; ${∆}_{s1}$ and ${∆}_{s2}$ represent the horizontal displacements on the left and right sides of the wall, respectively; $d$ denotes the diagonal length; $h$ is the wall section height; and ${\delta }_1-{\delta }_4$ represent the elongations from wire displacement transducers CP1–CP4, respectively.

Strain gauges were installed on the specimens to measure the strain in the steel sections and reinforcement bars. Their locations and quantities are depicted in Figure 6.

3. Experimental Results

3.1. Failure Process

Figure 7 illustrates the crack development of specimen RC-2.0-0.3-C50. Initially, no cracks were observed, but as the horizontal displacement reached 3.5 mm, transverse cracks began to emerge along the edges of the middle and lower sections of the wall. The horizontal cracks started to develop along the slope toward the lower portion of the wall when the horizontal displacement reached 7 mm. As displacement increased, existing cracks extended into the lower middle part of the wall, and new cracks appeared in the upper middle section. At a horizontal displacement of 31.5 mm (H/50), significant concrete spalling occurred at the bottom of the wall, causing a rapid decrease in the specimen’s load-bearing capacity and necessitating the termination of the test.

Figure 8 shows the crack development of specimen CFST-2.0-0.3-C50. Horizontal cracks appeared on the lower left and right sides of the web at 12 mm of horizontal displacement. While at a displacement of 15 mm, vertical cracks emerged in the middle of the web, extending from top to bottom. As displacement increased, diagonal and vertical cracks became more pronounced. Concrete began to spall at a horizontal displacement of 30 mm (approximately H/50), and the bottom of the steel tube bulged and buckled (Figure 8d). Tearing occurred in the bulged region of the steel tube at a displacement of 40 mm (approximately H/30) (Figure 8f).

The failure modes of the other specimens at the end of the examinations are depicted in Figure 9. CFST-2.0-0.2-C50 showed minor cracking at the base of the shear walls, as illustrated in Figure 9a,b, but extensive tearing at the bases of the steel tubes. CFST-2.0-0.4-C50 demonstrated a crack pattern similar to CFST-2.0-0.3-C50 but with more pronounced web damage and a higher density of cracks. The failure modes of the CFST specimens were significantly influenced by the restraining effect of the CFST frame on the web. With increasing axial compression ratio, the CFST frame exerted greater restraint on the web, leading to an increase in both the number and the severity of the web cracks, while the tensile load capacity of the CFST columns decreased. Figure 9c,d show that CFST-2.0-0.3-C40 and CFST-2.0-0.3-C50 exhibited similar crack patterns, with both specimens showing tearing at the bottom of the steel tube. CFST-2.0-0.3-C60 showed minimal web cracking along with severe buckling occurring at the base of the steel tube, without any tearing. This indicates that, although concrete strength increased, the corresponding stiffness of the web and CFST frame did not scale proportionally, resulting in inconsistent failure patterns across the specimens with varying concrete strengths. CFST-1.5-0.3-C50 showed severe web cracking, primarily composed of diagonal cracks, and buckling at the steel tube base, as demonstrated in Figure 9e,f. CFST-2.5-0.3-C50 displayed minimal cracking at the bottom of the web but showed signs of concrete crushing in that area. The observed finding indicates that the web of CRCSW-CFSTs experienced diagonal compression forces. Diagonal cracks were observed in regions with small vertical shear forces. In areas with moderate shear, vertical cracks appeared at the top, accompanied by flexural–shear diagonal cracks at the bottom. In areas with large vertical shear, pressure on the bottom corners resulted in concrete crushing.

3.2. Hysteretic Curves

Figure 10 illustrates the hysteresis curves for each specimen, showing essentially linear behavior at small displacement amplitudes for all specimens. As displacement amplitudes increased, the hysteresis curves evolved into spindle shapes, revealing distinct yielding characteristics. CRCSW-CFSTs showed a more pronounced pinching effect compared to RCSWs.

3.3. Skeleton Curves

Peak values at each displacement loading level were sequentially connected to form skeleton curves, as depicted in Figure 11. The characteristic points derived from these curves are summarized in

Table 4. Main parameters.

Specimen	Py (kN)	Δy (mm)	Pm (kN)	Δm (mm)	Pu (kN)	Δu (mm)	E (kN·m)	μ
RC-2.0-0.3-C50	397	10.8	440	17.3	374	32.2	120.1	2.97
CFST-2.0-0.3-C50	313	10.9	375	26.9	319	37.0	177.0	3.39
CFST-2.0-0.2-C50	256	9.8	308	29.8	262	37.0	114.4	3.77
CFST-2.0-0.4-C50	366	11.0	448	24.9	381	34.5	132.3	3.14
CFST-2.0-0.3-C40	280	11.0	343	27.5	292	33.7	128.6	3.06
CFST-2.0-0.3-C60	341	16.9	393	29.9	334	44.2	125.5	2.64
CFST-1.5-0.3-C50	544	11.1	638	23.9	542	27.4	60.1	2.47
CFST-2.5-0.3-C50	220	10.6	260	29.9	221	33.2	108.5	3.15

. The analysis of these curves and characteristic points led to the following conclusions:

(1)

The peak load for RC-2.0-0.3-C50 exceeded that of CFST-2.0-0.3-C50; however, the load-bearing capability of RC-2.0-0.3-C50 decreased rapidly after reaching its peak. CFST-2.0-0.3-C50 demonstrated a 55.4% increase in peak displacement, a 15.6% increase in ultimate displacement, and a 14.1% enhancement in displacement ductility coefficient upon comparison with RC-2.0-0.3-C50. This indicates that CRCSW-CFSTs provide superior ductility over traditional RCSWs due to the confining effect of the concrete-filled steel tube boundary columns on the web [42].

(2)

Among the three CRCSW-CFST specimens with different axial compression ratios, CFST-2.0-0.3-C50 and CFST-2.0-0.4-C50 demonstrated an increase of 21.8% and 45% in peak load, respectively, compared to CFST-2.0-0.2-C50. Despite the increase in peak load, CFST-2.0-0.3-C50 and CFST-2.0-0.4-C50 showed a decrease in peak displacement of 9.73% and 16.44% and a reduction in ductility coefficient of 10.1% and 16.7%, respectively. The observed results suggest that the load-bearing capacity of the CRCSW-CFST specimens increased while the deformation capacity decreased with an increasing axial compression ratio. This observation can be attributed to the higher axial pressure, which increased the shear compression zone height in the web. This enhancement in shear capacity reduced concrete cracking. However, in the specimens with higher axial pressures, the increased horizontal displacement accelerated concrete failure, leading to a rapid decrease in load-bearing capacity.

(3)

Among the three CRCSW-CFST specimens with varying concrete strengths, CFST-2.0-0.3-C50 and CFST-2.0-0.3-C60 demonstrated peak loads that were 9.33% and 14.58% higher, respectively, than that of CFST-2.0-0.3-C40. Furthermore, their peak displacements increased by 9.79 and 31.1%, respectively. The results indicate that increasing concrete strength enhances the load-bearing capacity of CRCSW-CFSTs.

(4)

Among the three CRCSW-CFST specimens with varying shear span ratios, CFST-2.0-0.3-C50 and CFST-2.5-0.3-C50 demonstrated 41.22 and 59.25% lower peak loads compared to CFST-1.5-0.3-C50. However, their peak displacements increased by 12.6% and 25.1%, respectively. The results showed that lower shear span ratios enhanced the load-bearing capacity but resulted in a more rapid decrease in capacity after reaching the peak load. The relationship between ductility and shear span ratio in CRCSW-CFSTs was not significant under the influence of diverse failure modes.

3.4. Energy Dissipation

The equivalent viscous damper coefficient (h_e) is a crucial indicator for assessing the energy dissipation capacity of a structure and can be calculated using Equation (4). The higher values of $h_e$ improved the energy dissipation performance of the shear walls [43].

$$h_e={\displaystyle \frac{1}{2\pi }}·{\displaystyle \frac{S_{(ABC+CDA)}}{S_{(OBE+ODF)}}}$$

(4)

where $S_{(ABC+CDA)}$ and $S_{(OBE+ODF)}$ represent the areas enclosed by the hysteresis curve and the triangle presented in Figure 12.

Figure 13 presents the $h_e$ values for eight specimens at various load levels, representing the average values from three cycles at each load level. The following conclusions can be drawn from the figure:

(1)

The $h_e$ value of RC-2.0-0.3-C50 was very low at small displacements (less than 15 mm), while that of CFST-2.0-0.3-C50 was relatively high. The $h_e$ value for RC-2.0-0.3-C50 increased rapidly after displacement exceeded 18 mm, whereas the value for CFST-2.0-0.3-C50 remained relatively stable. At 30 mm displacement, the $h_e$ value of CFST-2.0-0.3-C50 rose significantly, primarily because the $h_e$ values for RCSWs are mainly derived from plastic damage in the concrete and relative slips between the reinforcement and the concrete. At small displacement values, concrete plastic damage and reinforcement slip were minimal. However, as displacement increased, both concrete damage and reinforcement slip rapidly developed. For CRCSW-CFSTs, the slight relative slippage between the prefabricated concrete wall panel and the encasing C-shaped steel resulted in higher $h_e$ values at small displacements. Furthermore, the constraint provided by the concrete-filled steel tubes on the prefabricated web panels delayed plastic damage in the web, causing $h_e$ increase to begin only in the later stages of loading.

(2)

The $h_e$ value increased with the axial compression ratio among the three CRCSW-CFSTs, particularly at small displacements. This increase was primarily due to larger vertical loads, which made relative slippage between the reinforced concrete wall panel of the prefabricated web and the encasing C-shaped steel less likely to occur. At larger displacements, higher vertical loads resulted in more plastic damage to the concrete and greater relative slips between the reinforcement and the concrete.

(3)

The trends in the viscous damping coefficient for the three CRCSW-CFSTs with varying concrete strengths were similar. Both CFST-2.0-0.3-C40 and CFST-2.0-0.3-C50 showed comparable failure characteristics, leading to similar $h_e$ curves. In contrast, the failure of the CFST-2.0-0.3-C60 specimen was primarily due to the buckling and tearing of the bottom concrete-filled steel tube columns, which constrained the plastic energy dissipation capacity of the prefabricated web panels and resulted in the lowest $h_e$ value.

(4)

Among the three CRCSW-CFST specimens with different shear span ratios, those with higher shear span ratios exhibited higher $h_e$ values. This took place primarily because the relative slippage between the reinforced concrete wall panel of the prefabricated web and the encasing C-shaped steel in the specimens with higher shear span ratios consumed more energy. At larger displacements, the $h_e$ values showed varying trends due to the differing failure modes of the specimens.

3.5. Stiffness Degradation

The stiffness degradation curves for the specimens were derived by plotting the sinusoidal stiffness values $K_i$ at various amplitudes [44]. These curves were used to assess the degree of stiffness reduction in the specimens, which can be expressed as follows:

$$K_i={\displaystyle \frac{\left|+F_i\right|+\left|-F_i\right|}{\left|+X_i\right|+\left|-X_i\right|}}$$

(5)

where $K_i$ is the secant stiffness; $+F_i$ and $-F_i$ are the load values at the i-th positive and negative peak points, respectively; and $+X_i$ and $-X_i$ are the displacement values at the $i$ positive and negative peak points, respectively.

Figure 14 presents the stiffness degradation curves for all specimens. The figure illustrates that the stiffness degradation curves exhibited a nonlinear decrease with increasing displacement. Stiffness decreased rapidly during the initial loading phase and then tended to stabilize as displacement increased. The following conclusions were drawn from these observations:

(1)

The sinusoidal stiffness of the RC-2.0-0.3-C50 specimen was found to be higher than that of the CFST-2.0-0.3-C50 specimen during the early-to-mid stages of displacement loading. However, the stiffness of the RC-2.0-0.3-C50 specimen decreased more rapidly as displacement increased.

(2)

The stiffness degradation curves of CRCSW-CFSTs were largely consistent across different axial compression ratios and concrete strengths, indicating that these factors had minimal impact on the stiffness degradation of CRCSW-CFSTs.

(3)

The CFST-1.5-0.3-C50 specimen exhibited the highest positive stiffness during displacement loading but also experienced the most rapid decrease in rigidity as displacement increased. However, the CFST-2.5-0.3-C50 specimen displayed the opposite behavior, indicating that an increase in the shear span ratio improved the positive stiffness of CRCSW-CFSTs while slowing the rate of stiffness reduction with increasing load displacement.

3.6. Strain

Figure 15 displays the strain in each specimen in Section 1. Positive loading resulted in tension in the left CFST column and compression in the right CFST column at yield, peak, and ultimate loads for the CRCSW-CFST specimens. The figure reveals that the strains in the CRCSW-CFST steel tubes were significantly higher compared to those in the vertical distribution reinforcements of the web, primarily due to slippage between the concrete slabs of the prefabricated web and the C-shaped steel.

3.7. Deformation

The proportion of shear deformation to total deformation (α coefficient) was calculated using Equation (6).

$$\alpha ={\displaystyle \frac{{\gamma }_s}{{\gamma }_c+{\gamma }_s}}$$

(6)

Figure 16 shows the proportions of shear deformation to total deformation for each specimen at different displacements. The figure shows that the α coefficient increased with load displacement for both cast-in-place reinforced concrete and CRCSW-CFSTs. However, at peak load, the α coefficient for CRCSW-CFSTs with varying parameters was approximately 50%, whereas for cast-in-place reinforced concrete shear walls, it was around 33%.

4. Conclusions

In this study, a novel CRCSW-CFST frame was proposed, and its seismic performance was evaluated through experimental analysis. Furthermore, a method for calculating the shear-bearing capacity of CRCSW-CFSTs was developed based on the experimental findings. This study yielded the following conclusions:

(1)

The failure modes of CRCSW-CFSTs differ from those observed in RCSWs. The prefabricated web was subjected to diagonal compressive loads under cyclic loading, while the CFST frame columns experienced bending and shear forces. The final failure mode was influenced by factors such as the axial compression ratio, concrete strength, and shear span ratio, leading to three main scenarios: 1. Tearing in the low-plasticity region of the CFST frame columns, with minimal or no cracking in the prefabricated web. 2. Tearing in the low-plasticity region of the CFST frame columns, accompanied by severe cracking and crushing in the prefabricated web. 3. Minor bulging at the base of the CFST frame columns, along with cracking and significant crushing in the prefabricated web.

(2)

The weld seam remained intact throughout the test, demonstrating the reliability of the welded connection.

(3)

As the axial compression ratio increased, crack development in the prefabricated web increased, while damage to the CFST frame columns decreased. Increasing the axial compression ratio from 0.2 to 0.3 and 0.4 raised the peak load of CRCSW-CFSTs by 21.8% and 45%, respectively, but reduced the peak displacement by 9.73% and 16.44%. Additionally, as the axial compression ratio increased, the stiffness before peak load improved, while both the ultimate displacement and ductility coefficient declined. Moreover, strength degradation accelerated after exceeding peak load.

(4)

An increase in concrete strength from C30 to C40 and C50 raised the peak loads of CRCSW-CFSTs by 9.33% and 14.58%, respectively, and enhanced the ultimate displacements by 9.79% and 31.1%. Meanwhile, the yield capacity and stiffness remained relatively unchanged.

(5)

Increasing the shear span ratio from 1.5 to 2.0 and 2.5 reduced the yield load of CRCSW-CFSTs by 41.22% and 59.25%, respectively, while the peak displacement increased by 12.6% and 25.1%, respectively. Additionally, the stiffness of CRCSW-CFSTs decreased with a higher shear-to-span ratio, though the rate of stiffness reduction slowed as displacement increased.

(6)

In the initial loading stages, the energy dissipation capacity of the shear wall primarily resulted from minor relative slippage between the prefabricated reinforced concrete panel and the encasing C-shaped steel.