Experimental Lateral Behavior of Porcelain-Clad Cold-Formed Steel Shear Walls Under Cyclic-Gravity Loading

Abstract

Lightweight steel-framing (LSF) systems have become increasingly prominent in modern construction due to their structural efficiency, design flexibility, and sustainability. However, traditional facade materials such as stone are often cost-prohibitive, and brick veneers—despite their popularity—pose seismic performance concerns. This study introduces an innovative porcelain sheathing system for cold-formed steel (CFS) shear walls. Porcelain has no veins thus it offers integrated and reliable strength unlike granite. Four full-scale CFS shear walls incorporating screwed porcelain sheathing (SPS) were tested under combined cyclic lateral and constant gravity loading. The experimental program investigated key seismic characteristics, including lateral stiffness and strength, deformation capacity, failure modes, and energy dissipation, to calculate the system response modification factor (R). The test results showed that configurations with horizontal sheathing, double mid-studs, and three blocking rows improved performance, achieving up to 21.1 kN lateral resistance and 2.5% drift capacity. The average R-factor was 4.2, which exceeds the current design code values (AISI S213: R = 3; AS/NZS 4600: R = 2), suggesting the enhanced seismic resilience of the SPS-CFS system. This study also proposes design improvements to reduce the risk of brittle failure and enhance inelastic behavior. In addition, the results inform discussions on permissible building heights and contribute to the advancement of CFS design codes for seismic regions.

1. Introduction

Cold-formed steel (CFS) systems have revolutionized low-rise building construction owing to their lightweight properties and ease of assembly [1]. Many innovative CFS sheathed walls with advantages over conventional configurations have been discussed in the literature [2]. Despite this progress, seismic design methodologies for CFS systems remain underdeveloped, and current seismic codes lack comprehensive guidelines for these systems [3,4,5]. Therefore, further research is essential to enhance the knowledge of key seismic parameters for CFS shear walls, including response modification factors (R-factors), load-bearing capacity, and ductility. Extensive research has emphasized the critical role of sheathing materials in influencing the lateral performance of CFS shear walls. Gad et al. [6] identified sheathing as a pivotal element affecting the behavior of lightweight steel frames (LSFs). The American Iron and Steel Institute (AISI) has certified materials like steel, gypsum, wood, and concrete—typically secured with self-drilling screws—for use in CFS-LSFs [7]. Masonry infills, such as bricks or stones bonded with adhesives, have also been employed; however, challenges in material compatibility have driven a shift toward LSF-based systems. While many studies have analyzed the structural components of these materials, most have concentrated on wet-set applications, which involve labor-intensive facade cladding processes. Esmaili et al. [8], for example, investigated mortar dimensions in brick-sheathed LSFs. Although Sharafi et al. [2] and Usefi et al. [9] synthesized findings from such studies, the influence of screwed porcelain sheathing (SPS) on the seismic resilience of CFS shear walls has not yet been thoroughly examined.

AISI S400 [10] formally acknowledges CFS shear walls with steel or wood sheathing but excludes gypsum-based systems. Nonetheless, materials like gypsum wall boards (GWBs), Bolivian magnesium boards (BMBs), and calcium silicate boards (CSBs) are commonly utilized for both fire resistance and lateral bracing of studs [11]. Many standards, such as those in Iran [12], allow the use of these materials in lateral load-resisting systems, provided that they meet the specified performance criteria.

Porcelain panels offer a unique combination of properties that are highly advantageous for earthquake-resistant construction: low density, high compressive strength, resistance to cracking, and compatibility with mechanical dry fastening systems. Dry-set porcelain cladding systems fastened with screws eliminate adhesive curing time, promote faster on-site assembly, and reduce the total building mass. In contrast, wet-set veneers like masonry tiles or bricks, are not only labor-intensive but also introduce substantial additional weight, which is undesirable in seismic zones. Despite these potential benefits, no comprehensive experimental study has assessed the seismic performance of the SPS-CFS system under cyclic lateral and gravity loadings.

The high construction speed (achieved by replacing wet construction with dry construction through the use of screw connections attaching the cladding to LSF, eliminating the need for infill walls) and the removal of the requirement for separate labor-intensive facade cladding processes are among the advantages of innovative SPS-CFS systems; hence, their study is undoubtedly valuable. The present study addresses this significant knowledge gap by conducting full-scale cyclic and gravity-load tests on four SPS-CFS wall specimens. This study investigated key structural parameters, including sheathing orientation (horizontal vs. vertical), mid-stud configuration (single vs. double), and the presence or absence of intermediate blocking rows. The lateral strength, drift capacity, energy dissipation, and effective response modification factor (R) were evaluated and compared with both conventional systems and existing code limits. This research builds upon the foundational contributions of Yu et al. [13] and Li et al. [14], extending the design space into previously unexplored ceramic-based sheathing domains.

Further reinforcement for this research comes from recent studies. Pan and Shan [15] investigated the performance of GWB, CSB, and oriented strand board (OSB) sheathing under monotonic shear conditions. Their study identified screw-induced sheath tearing and panel-screw separation as the primary failure mechanisms. The OSB-sheathed frames demonstrated the maximum shear resistance, whereas the GWB specimens demonstrated the lowest. Walls with a 2:1 aspect ratio exhibited a 35% reduction in shear capacity compared with 1:1 configurations. The CSB systems displayed superior energy absorption, while the GWB performed poorly; the ductility trends correlated with the screw arrangement and bracing design. The single-sided sheathing outperformed the double-sided setup in terms of ductility. Baran and Alica [16] observed that hold-down fixtures in CFS wall tracks markedly influenced the structural behavior, whereas diagonal bracing minimally affected the load-bearing capacity and initial stiffness. Swensen et al. [17] noted that contemporary screw designs and adhesive-based sheathing-frame bonding substantially enhanced connection strength over traditional methods. Mohebbi et al. [18] evaluated steel-sheathed CFS walls under cyclic loads, revealing that walls with failed sheathing connections absorbed more energy than those with buckling studs. Double-sided sheathing improved energy dissipation, strength, and stiffness when stud failure was mitigated. Niari et al. [19] determined that doubling the steel sheathing thickness increased the shear capacity by 42%, with failure modes tied to sheathing-fastener separation. Javaheri et al. [20] tested steel-sheathed walls with varied screw layouts, advocating an increased response modification factor from 6.5 to 7. Closer screw spacing (150 to 100 mm) improved the shear strength in single-end stud designs by 16–18%, although double-end configurations showed no gains. Their study involved 24 full-scale walls in eight configurations applied to three 1.2 × 2.4-m panels. Zhang et al. [21] analyzed corrugated steel-sheathed CFS walls under gravity loads, noting that gravitational loads at the service load level enhanced stiffness and strength, with a proposed 7% drift limit for collapse. Zeynalian and Ronagh [22] assessed fiber cement board (FCB)-sheathed systems, deeming their seismic performance inadequate (R = 2.5) but improved to R = 5 when paired with X-bracing. Xu et al. [23,24] introduced high-strength lightweight foamed concrete (HLFC) sheathing, which increased lateral strength while shifting failure modes from brittle to ductile. Higher foam grades and stud thickness enhanced the performance, although vertical loads reduced the ductility. Wu and Li [25] conducted multi-level numerical simulations on steel-sheathed CFS wall systems and proposed a higher R-factor value (R = 4.25), significantly beyond current conservative code estimates. Liu et al. [26] used numerical simulations to show that combining steel with gypsum board sheathing improved both lateral stiffness and ductility under dynamic excitation. Lopes et al. [27] experimentally validated similar enhancements in LSF systems, especially when attention was paid to the screw spacing and sheathing-to-frame connection details. However, these studies predominantly focused on conventional materials such as OSB, GWB, and fiber cement board. Alternative sheathing materials—such as ceramics and specifically porcelain—remain unexplored in the context of seismic loading. Iuorio [28] offered a broader perspective by discussing the role of lightweight prefabricated CFS systems in mass customization and seismic design. His findings emphasized how dry-set ceramic panels could revolutionize seismic construction by combining efficiency, reduced weight, and high crack resistance in an industrialized framework.

The AISI standards [29,30] outline the criteria for CFS shear walls with R-factors below 3, emphasizing uplift resistance and boundary elements. However, they exclude newer systems like SPS-CFS walls. FEMA 450 [31] and FEMA P750 [32] under the National Earthquake Hazard Reduction Program (NEHRP) mandate AISI compliance but face inconsistencies in the load/stress criteria, requiring calibration. Many CFS systems lack approved R-factors and default to 3. AISI S400 [10] recommends R = 4 for diagonally strapped CFS walls. The Australian/New Zealand Standard (AS/NZS 4600) [33] restricts R-factors to 2 in Australia/New Zealand, where seismic risks are low and wind governs the design.

A simple but important conclusion from the above review is that there is no universal agreement on the value of the response modification factor, R, and in particular, there is no reference in the codes for the R-factor of the SPS-CFS system. Therefore, further studies are required to clarify this issue.

In light of recent developments, the present investigation pioneers the full-scale evaluation of SPS-CFS shear wall systems, with a special focus on cyclic degradation, R-factor quantification, and performance comparison with both conventional systems and international seismic design standards. The findings of this study offer critical insights for structural engineers and code developers aiming to integrate advanced sheathing solutions like porcelain, into mainstream seismic construction practices.

2. Experimental Description

2.1. Program of Experimental

Four full-scale CFS shear wall specimens, each measuring 1.2 × 2.4 m [20,34], were tested to assess their hysteretic response to cyclic lateral loading. The specimens represented different wall configurations to examine the influence of key design parameters, as listed in

Table 1. Specimen configurations.

Wall Panel	Dimensions (cm)	Porcelain Size (cm)	Chord Studs	Middle Studs	Stud Spacing (cm)	Sheathing Orientation	Blocking Rows	G * (kN)	Schematic
S-H-3	240 × 120	60 × 120	Double	Single	60	Horizontal	3	40	Figure 1A
D-H-3	240 × 120	60 × 120	Double	Double	60	Horizontal	3	40	Figure 1B
D-V-3	240 × 120	60 × 120	Double	Double	60	Vertical	3	40	Figure 1C
D-V-1	240 × 120	60 × 120	Double	Double	60	Vertical	1	40	Figure 1D

* Gravity load.

. The CFS framing members had a nominal thickness of 0.7 mm, and their mechanical properties are listed in

Table 2. The mechanical properties of the CFS sections.

Mechanical properties	Value	Mechanical properties	Value
Nominal grade	550 MPa	Yield strain	0.0045
Nominal thickness	0.7 mm	Ultimate stress, Fu	617 MPa
Elastic modulus	169 GPa	Ultimate strain	0.0286
Yield stress, Fy	592 MPa	Fu/Fy	1.04

. Each specimen was sheathed on one side with porcelain panels and on the other side with plasterboard. The sheathing was attached using self-drilling screws with a minimum embedment of three thread pitches to ensure strong connections. The fabrication and testing were conducted at the Structural and Earthquake Research Center, Islamic Azad University, Taft Branch.

Porcelain was chosen due to its superior strength-to-weight ratio and suitability for dry construction methods compared with granite, along with better crack resistance and higher durability. The porcelain panels used in this study were 11 mm thick, with a tensile strength of 8.5 MPa, compressive strength of 130 MPa, and density of 2200 kg/m³.

The specimen labels were designed to reflect the key configuration variables:

• S/D = Single or double middle studs
• H/V = Horizontal or vertical porcelain sheathing orientation
• Number = Number of blocking rows

For example, specimen D-V-3 denotes a wall with double middle studs, vertical sheathing, and three blocking rows.

Three configuration comparisons were used to isolate the specific effects:

• S-H-3 vs. D-H-3 → Effect of mid-stud quantity
• D-H-3 vs. D-V-3 → Effect of sheathing orientation
• D-V-3 vs. D-V-1 → Effect of number of blocking rows

This parametric approach enabled a systematic assessment of the seismic performance parameters under controlled cyclic loading.

Figure 1 shows the layout of the porcelain sheathing, including the panel dimensions and placement. To avoid premature separation during failure, the joints between the frame and porcelain were sealed using silicone. In addition, to reduce stress concentrations in the brittle porcelain material, the screw spacing was set at 15 cm along the panel edges and 20 cm in the interior zones. Figure 2 and Figure 3 provide additional details.

2.2. Test Setup and Instrumentation

A reliable experimental configuration is essential for capturing accurate seismic behavior data. Figure 4 depicts a schematic of the testing apparatus and support system engineered to accommodate specimens as large as 2.4 × 2.4 m. The specimens were secured between a stationary upper support beam and a rigid lower loading beam using high-strength M16 bolts. These bolts were torqued to roughly 190 Nm (corresponding to 53 kN of tension) to establish a stable connection and eliminate slippage risks. Hold-down devices were installed at each corner of the wall to mitigate overturning tendencies and maintain effective load transfer from the bracing elements to the wall chords and studs [20,34]. Figure 5 shows the configuration of the gusset angles and associated bolts.

A linear variable displacement transducer (LVDT) was deployed to record the horizontal displacements at the bottom track, enabling precise monitoring of lateral shifts during cyclic loading. Simultaneously, a load cell measured the racking resistance, with all data transmitted to a computerized system for the real-time generation of load-displacement curves. Vertical gravity loads were applied using two 200 kN hydraulic jacks. A flow rail roller positioned between the loading beam and jacks ensured that the loading point was synchronized with the wall panel’s lateral displacement. Notably, the gravity load remained constant throughout the testing. This integrated instrumentation framework enables a thorough evaluation of the structural response to seismic forces.

2.3. Lateral and Gravitational Loading Protocol

Cyclic loading followed Method B of ASTM E2126 [35], in which the specimens were subjected to displacement-controlled cycles until failure or substantial loss of load capacity.

Table 3. Cyclic displacement schedule.

Loading Step	1	2	3	4	5	6	7	8
ASTM displacement (% of ${∆}_{\mathrm{u}}$)	1.25	2.5	5	7.5	10	20	40	60
Current study displacement (% of ${∆}_{\mathrm{t}}$)	2.5	5	10	15	20	40	80	120 *
Displacement amplitude (mm)	1.5	3	6	9	12	24	48	72 *

* Note: Due to actuator travel limitations, the 8th step was capped at 60 mm instead of 72 mm.

and Figure 6 summarize the displacement protocols.

The actuator had a maximum displacement of 75 mm, equivalent to a 3.125% inter-story drift, exceeding FEMA 450’s limit of 2.5% and ensuring a conservative safety margin. The loading rate was maintained constant at 2 mm/s in accordance with the ASTM E2126-07 guidelines to ensure repeatability.

Before applying the lateral loads, an axial force of 40 kN was applied to the bottom track of each specimen. This value represents the axial capacity of the bare CFS frame prior to local buckling, as determined through preliminary incremental compression tests. This also corresponded to the allowable compressive strength of each frame, calculated as 0.6 × P_n × (number of studs), where P_n denotes the nominal compressive strength of a thin-walled stud section [12,36]. The gravity loading remained constant throughout the test to simulate the axial force conditions in multi-story buildings.

3. Evaluation of Seismic Response Modification Factor

To achieve a balance between safety and economy, modern seismic design seeks to reduce seismic demands through ductile performance and energy dissipation. This concept is embodied in the response modification factor (R), which combines two components: the ductility reduction factor (R_d) and the structural over-strength factor (${\mathsf{\Omega }}_0$) [31,32,37]. This is mathematically expressed as

$$R=R_d\times {\Omega }_0$$

(1)

as depicted in Figure 7, the factors are defined as follows:

$$R_d={\displaystyle \frac{V_e}{V_y}}$$

(2)

$${\Omega }_0={\displaystyle \frac{V_y}{V_s}}$$

(3)

thus, the R-factor can also be derived as:

$$R={\displaystyle \frac{V_e}{V_s}}$$

(4)

here, ${\mathrm{V}}_{\mathrm{e}}$ represents the elastic response strength, ${\mathrm{V}}_{\mathrm{y}}$ is the idealized yield strength, and ${\mathrm{V}}_{\mathrm{s}}$ denotes the first significant yield strength.

Figure 7 illustrates how these factors are derived using the actual load-displacement behavior, equivalent elastic curve, and idealized bilinear response. The idealized yield strength (${\mathrm{V}}_{\mathrm{y}}$), the actual elastic response strength (${\mathrm{V}}_{\mathrm{e}}$), and the design strength (${\mathrm{V}}_{\mathrm{s}}$) were determined graphically.

According to FEMA 356 [38], the idealized curve is formed by two lines determined through the equality of the area under the actual curve and the area under the idealized curve using an iterative graphical method. The first line passes the intersection of the origin and the point on the actual curve at 0.6 ${\mathrm{V}}_{\mathrm{y}}$, whereas the second line connects the first line at ${\mathrm{V}}_{\mathrm{y}}$ to the target displacement (${∆}_{\mathrm{t}}$), defined as the maximum drift before a substantial strength reduction. ${\mathrm{V}}_{\mathrm{y}}$, must not exceed the actual curve’s peak base shear. For this study, ${∆}_{\mathrm{t}}$ was taken as 60 mm, consistent with FEMA 450 [31] guidelines that define allowable story drift as 0.025 H, or earlier displacement corresponding to a 20% drop in strength.

The R_d factor reflects a structure’s inelastic deformation capacity, which is influenced by ductility, energy dissipation, fundamental period, and soil properties [39,40]. Newmark and Hall [41] established widely used relationships between R_d and ductility (μ) as follows: Structure ductility is defined by Equation (5). Where, ${∆}_{\mathrm{m}\mathrm{a}\mathrm{x}}$ is the maximum drift and ${∆}_{\mathrm{y}}$ is the idealized yield drift.

$$\mu ={\displaystyle \frac{{\mathsf{\Delta }}_{max}}{{\mathsf{\Delta }}_y}}$$

(5)

For varying fundamental periods (T), the R_d-factor can be calculated using the following sub-equations which have been widely adopted in the literature [34,40]:

$$\mathrm{Newmark} \& \mathrm{Hall}\text{:} R_d=\left\{\begin{array}{c}\mu T > 0.5 \mathrm{S}\\ \sqrt{2\mu -1} 0.1<T< 0.5 \mathrm{S}\\ 1 T < 0.03 \mathrm{S}\end{array}\right.$$

(6)

Therefore, the R_d-factor can be derived either geometrically through Figure 7 or using the Newmark-Hall method, and a good agreement exists between the two methods [34]. Hence, Equation (6) was used to assess R_d (sub-equation with T = 0.1–0.5 s for LSF structures [6,34]). ${\mathrm{V}}_{\mathrm{s}}$ (design strength) was determined by back-calculating design capacities from tested specimens at first yield [20]. To refine the determination of ${\mathrm{V}}_{\mathrm{s}}$, this study defined it as the base shear at which the actual response curve deviated by 0.005 ${∆}_{\mathrm{t}}$ from the idealized bilinear curve along the displacement axis. The corresponding displacement at this point was designated as ${∆}_{\mathrm{s}}$, and ${\mathsf{\Omega }}_0$ was derived via Equation (3).

4. Test Results and Discussion

4.1. Failure Modes

Visual inspection during testing showed that the initial failure typically occurred as cracking in the plasterboard corners at a displacement of 24 mm, corresponding to the sixth loading step. In specimen D-H-3, additional shear failure in the screw connections of the porcelain sheathing was observed near the corner of the frame.

Table 4. Failure observations during lateral loading.

Specimen	Failure Onset Displacement, ∆sf (mm)			Fractures Count			${∆}_{\mathbf{t}}$ (mm)	$\mathbf{Drift} \mathbf{Ratio}, {∆}_{\mathbf{t}}$/H	$\mathbf{Max}. \mathbf{Strength}, {\mathbf{V}}_{\mathbf{m}\mathbf{a}\mathbf{x}}$ (kN)
	Board	Porcelain	Screw	Board	Porcelain	Total
S-H-3	24	48	48	4	4	8	58	0.024	20
D-H-3	24	48	24	4	2	6	60	0.025	21.1
D-V-3	24	48	48	4	3	7	48	0.02	17.4
D-V-1	24	48	60	4	3	7	51	0.021	19.8

summarizes the failure onset, location, and severity across all specimens. Ceramic cracking was consistently observed at a displacement of 48 mm.

These findings suggest that the presence of plasterboard on the opposite side of the porcelain sheathing may accelerate wall degradation by introducing differential deformation behavior. The representative damage patterns are shown in Figure 8.

The severity of the observed damage was generally similar across the specimens, although it was slightly more pronounced in S-H-3, likely due to the use of a single middle stud. While plasterboard cracking appeared relatively insensitive to specimen configuration, porcelain damage decreased by over 25% when double middle studs were employed instead of a single stud.

Failures in the screw connections of the porcelain panels (Table 4) were primarily influenced by three variables: the number of middle studs, orientation of the sheathing strips, and number of blocking rows. In some cases, the screw heads fractured under cyclic loading; however, no loss of overall structural integrity was observed, indicating adequate redundancy in the fastener design. Among all configurations, specimen D-H-3 achieved the highest ultimate lateral strength (21.1 kN) and drift capacity (2.5%).

4.2. Hysteresis Response and Envelope Curves

Hysteresis loops and envelope curves were generated from the recorded load–displacement data (Figure 9 and Figure 10) to evaluate the nonlinear cyclic response of each wall specimen. Key seismic parameters—including lateral strength (V), ductility (μ), stiffness (K) and energy dissipation (E)—were extracted from these curves.

The hysteresis curves of the specimens were highly similar. Among the tested configurations, D-H-3 exhibited better centering and more stable hysteresis loops. The hysteresis curves exhibited an elastic behavior during the initial loading phase. Up to a displacement of 24 mm, no failures were observed, and the shear force at this displacement was close to 15 kN. Upon reaching a displacement of 48 mm and the appearance of cracks in the porcelain sheathing, the shear resistance nearly reached its ultimate value, the hysteresis loops widened, and the pinching phenomenon was observed.

4.3. Energy Dissipation

The energy dissipation of the test specimens was determined by calculating the area of the hysteresis loop [42,43]. The energy dissipation per loading step (

Table 5. Energy dissipation per loading step (Joules).

	Specimen	S-H-3	D-H-3	D-V-3	D-V-1	Average
Step
1		9	9	8	9	9
2		27	27	25	28	27
3		70	67	62	73	68
4		147	129	125	142	136
5		251	217	220	244	233
6		1180	963	834	1048	1006
7		3436	3075	2829	3125	3116
8		3538	5503	2829	3625	3874
Total:		8658	9990	6932	8294	8469

) was computed by summing the areas enclosed within the corresponding hysteresis loops.

Notable energy dissipation began at Step 6 (Figure 11), following the initial yielding observed at the end of Step 5, corresponding to a displacement of 12 mm. This indicates that all specimens behaved elastically up to a drift ratio of 0.005 H, which corresponds to the typical operational earthquake limits. Therefore, the base shear at this threshold may be considered the design-level strength.

4.4. Stiffness and Other Seismic Characteristics

Figure 12 illustrates the actual and idealized structural responses. The idealized curves were drawn, and the R-factors were calculated using the procedures outlined in Section 3 (Evaluation of Seismic Response Modification Factor) and are presented in

Table 6. Response modification factors.

Specimen	Direction	Vy (kN)	∆y (mm)	∆t (mm)	Vs (kN)	∆s (mm)	µ	Rd	Ω0	R	Cd (∆t/∆s)	Ki (Vy/∆y)	Vmax (kN)
	Positive	12.9	14.2	60	8.1	9.2	4.2	2.72	1.59	4.3	6.52	0.9	18.8
S-H-3	Average		16							4
	Negative	16.38	17.74	56.4	10.28	11.4	3.18	2.3	1.59	3.7	4.95	0.9	21.2
	Positive	17.11	26.8	60	10.6	16.88	2.2	1.9	1.61	3.1	3.55	0.6	21.5
D-H-3	Average		19							4.1
	Negative	12.05	10.67	60	7.53	6.95	5.6	3.2	1.6	5.1	8.63	1.1	20.7
	Positive	12.6	11	48	7.87	7.14	4.4	2.8	1.6	4.5	6.72	1.1	16.9
D-V-3	Average		13							4.1
	Negative	12	14.4	48	7.7	9.5	3.33	2.4	1.56	3.7	5.05	0.8	17.9
	Positive	18.2	22.12	54	11.2	13.8	2.44	1.97	1.63	3.2	3.91	0.8	22
D-V-1	Average		14							4.4
	Negative	8	6.7	48	5.2	4.6	7.16	3.65	1.54	5.6	10.43	1.2	17.6

.

Table 7. Key seismic characteristics of specimens.

Specimen	Vy (kN)	Vmax (kN)	E * (Joule)	∆t/H (%)	Ki (kN/mm)	Cd	µ	Ω0	R
S-H-3	14.6	20	8658	2.4	0.9	5.7	3.7	1.6	4
D-H-3	14.6	21.1	9990	2.5	0.85	6.1	3.9	1.6	4.1
D-V-3	12.1	17.4	6932	2	0.95	5.9	3.9	1.6	4.1
D-V-1	13.3	19.8	8294	2.1	1	7.2	4.8	1.6	4.4
Average	13.7	19.6	8469	2.3	0.9	6.2	4.1	1.6	4.2

* Cumulative energy dissipation.

summarizes the initial stiffness (K_i), dissipated energy (E), ductility factor, ultimate strength (V_max), lateral displacement magnification ratio (C_d), and other key seismic parameters.

Using the idealized curve, the initial stiffness of the specimens (K_i) was calculated as the force ratio to the displacement of the elastic region, as shown in Table 6. The stiffness degradation of the test specimens was described using secant stiffness [42]. The secant stiffness equals the gradient of the line passing between (0,0) and point ith (∆_i, V_i) on the actual curve. As the displacement increases, the specimen stiffness degrades, as shown in Figure 13.

The computed R-factors for the SPS-CFS wall specimens ranged from 4.0 to 4.4, with an average of 4.2. These values indicate that the R = 3 value currently recommended by the AISI S213 and FEMA provisions is conservative for this system, while the AS/NZS 4600 value of R = 2 appears excessively restrictive. Interestingly, specimen D-V-3 exhibited the lowest ultimate strength and drift ratio, despite not having the lowest R-factor, as shown in Table 7.

Specimens with horizontally oriented porcelain panels demonstrated higher ultimate lateral strengths than those with vertically oriented panels (Table 7). As shown in Figure 12, D-H-3 maintained its stiffness and strength up to a full test displacement of 60 mm (equivalent to 0.025 H drift). It also achieved maximum strength, drift capacity, and energy dissipation, confirming D-H-3 as the most seismically efficient configuration of the SPS-CFS system.

The superior performance of the double-middle-stud configuration with horizontal porcelain sheathing can be attributed to its enhanced stiffness and better stress distribution. The horizontal arrangement helps distribute the loads more uniformly along the wall tracks and through the blocking elements, thereby reducing localized stress concentrations. In contrast, vertical sheathing tends to concentrate forces at the corners, making the system more prone to early cracking and failure.

Figure 14, Figure 15, Figure 16 and Figure 17 present the correlations among the key seismic performance indicators. A strong linear relationship was observed between the response modification factor (R) and displacement amplification factor (C_d), as shown in Figure 14 and Figure 15. Based on data from Table 7 and

Table 8. The ratio of C_d to R and related parameters.

Specimen	Cd	Cd/R	Ω0.R
S-H-3	5.7	1.4	6.4
D-H-3	6.1	1.5	6.6
D-V-3	5.9	1.4	6.6
D-V-1	7.2	1.6	7
Average	6.2	1.5	6.7

, the average C_d/R ratio was approximately 1.5, which closely corresponds to 0.94 times the average over-strength factor (${\mathsf{\Omega }}_0$), consistent with theoretical expectations (see Equation (7)).

$$C_d=0.94{\Omega }_0.R=1.5R$$

(7)

According to seismic design codes, the maximum elastic displacement (∆_e) at the service level must not exceed 0.005 H, where H is the story height. To avoid structural damage, the first-yield displacement (∆_s) must exceed this threshold. In this study, ∆_s was determined to be 12 mm, implying a maximum allowable wall height (H_max) of 2.4 m, under the current conditions.

However, if premature failure is addressed through reinforcements—such as employing CFS sections thicker than 0.7 mm, high-strength screw connections, and replacing plasterboard with porcelain on both sides—the critical displacement capacity can be raised to ∆_sf = 48 mm, thereby increasing H_max to 9.6 m (i.e., 48 mm/0.005).

For comparison, wall systems sheathed with conventional materials like GWB or fiber cement board typically achieve lower permissible heights; for instance, Iranian Standard 612 [12] limits the maximum height to 7.2 m under similar conditions.

5. Conclusions and Recommendations

In this article, an experimental study was conducted to investigate the performance of SPS-CFS shear walls. Four SPS-CFS shear wall specimens were constructed and examined under cyclic loadings. The use of porcelain sheathing in lightweight steel-framing (LSF) systems has demonstrated significant improvements in seismic behavior. The following conclusions and design recommendations are drawn based on the experimental findings:

Effect of middle stud quantity:

A comparison between specimens S-H-3 and D-H-3 indicated that adding a second middle stud improved the R-factor from 4.0 to 4.1 and slightly enhanced ductility (μ increased from 3.7 to 3.9), suggesting a marginal but measurable benefit to seismic capacity.

Effect of sheathing orientation:

A comparison between specimens D-H-3 and D-V-3 showed that the sheathing strip orientation (horizontal vs. vertical) had no significant impact on the R-factor or ductility, with both configurations yielding R ≈ 4.1 and μ ≈ 3.9.

Effect of blocking quantity:

Increasing the number of blocking rows from one (D-V-1) to three (D-V-3) led to a reduction in both the R-factor (from 4.4 to 4.1) and ductility (from 4.8 to 3.9), suggesting that excessive blocking may limit the deformation capacity.

Average system performance:

The SPS-CFS wall system exhibited an average R-factor of 4.2 and an ultimate base shear strength of 19.6 kN. These values demonstrate that the current code recommendations (R = 2–3) may be overly conservative, especially for horizontally sheathed systems with minimal blocking.

Relationship between R and C_d:

A strong linear relationship was identified between the displacement amplification factor (C_d) and R factor, with C_d ≈ 1.5 R. This empirical correlation supports the simplified design assumptions in performance-based seismic engineering.

Design implications and height limits:

SPS-CFS walls with porcelain sheathing may be limited to 2.4 m in height under the current configurations to remain within safe elastic response limits. However, through reinforcement measures—such as using thicker CFS sections, high-strength fasteners, and replacing plasterboard with porcelain on both sides—the allowable wall height may be increased to 9.6 m, exceeding existing code limits (e.g., 7.2 m per Iranian Standard 612).

The present findings were derived from controlled laboratory experiments, which may not comprehensively reflect the complexities of in situ conditions. Variability in material properties and construction methodologies can lead to deviations in the actual performance. Accordingly, future research should focus on (1) assessing the long-term durability and functional performance under a range of environmental exposures, (2) exploring alternative sheath materials and design configurations, and (3) validating laboratory outcomes through field-scale studies and real-world implementations to enhance the robustness and generalizability of the results.