Research Article  Open Access
Zhi Zhou, Jiang Qian, Wei Huang, "Numerical Study on Deformation Capacity of Steel Plate Reinforced Concrete Shear Walls", Advances in Civil Engineering, vol. 2019, Article ID 9701324, 13 pages, 2019. https://doi.org/10.1155/2019/9701324
Numerical Study on Deformation Capacity of Steel Plate Reinforced Concrete Shear Walls
Abstract
Steel plate reinforced concrete (SPRC) shear wall consists of steel plate encased in the concrete, in which the material advantages of both concrete and steel are utilized. The lateral resistance and deformation capacity of the shear wall are greatly improved. This paper investigates the deformation capacity of the SPRC shear wall under cyclic loads. A nonlinear 3D finite element model in ABAQUS was developed and validated against published experimental results. Then, a parametric study was conducted to obtain the yield and ultimate rotation of SPRC shear walls with flexural failure. By statistical analyses, formulas for the yield and ultimate rotation of SPRC shear wall were proposed.
1. Introduction
Steel plate reinforced concrete (SPRC) shear wall consists of steel plate encased in the concrete wall as shown in Figure 1 [1]. Concrete on two sides of steel plate can prevent it from local buckling, and the shear yielding mechanism of steel plate can result in more stable hysteretic loops for shear wall [2]. This type of combination of concrete and steel plate can be utilized to change the potential brittle shear failure of RC shear wall to a ductile failure, while the stiffness and strength of the shear wall can also be significantly improved. Therefore, SPRC composite walls are generally used in cases where shear walls are designed to resist extreme lateral loads under earthquakes.
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Experimental studies have been conducted by many researchers. Zhao and AstanehAsl [3] tested a series of composite shear walls and found them rather promising with a better capacity of energy dissipation. Gan et al. [4] conducted cyclic tests of 16 SPRC shear walls and investigated the effects of shear span ratio, the thickness of concrete and steel plate, and shear connectors between concrete and steel plate on the seismic behavior of shear walls. It shows that thicker concrete can provide more confinement to the steel plate and yield a more ductile failure and the shear studs can enhance the strength and ductility of SPRC shear wall. According to Wu [5], experimental results suggest that the increase of the steel plate ratio can significantly improve the strength of the SPRC shear wall but has a limited effect on its ductility. The strength and ductility of the SPRC shear wall are affected by the concrete thickness and boundary members. Jiang [6] tested 9 composite shear walls with different configurations under high axial load. It was concluded that the encased steel plate can greatly improve the flexural capacity of the shear wall and its deformation capacity degrades when designed axial load ratio exceeds 0.5. Zhu [7] tested 8 SPRC shear walls and 2 steel reinforced shear walls with a shear span ratio of 2.0. The SPRC shear walls were found to have a better capacity of energy dissipation than steelreinforced concrete shear wall. And the energy dissipation capacity of SPRC shear walls can be improved by either increasing the reinforcement ratio or the rebar space of the wall web. It was pointed out that the steel plate and concrete can deform coordinately with the shear studs and tie bars in the test. Based on cyclic tests of 16 composite shear walls, Wang et al. [8] proposed a hysteretic model for SPRC shear walls.
Numerical modelling has been carried out by researches to better understand the behavior of SPRC shear walls. Rahnavard et al. [9] conducted numerical analysis on different types of composite steelconcrete shear walls to study their hysteretic response, frame drift, and dissipated energy in Abaqus. The models involve concrete with and without gap to steel frame on one side and two sides of steel plate. It was concluded that the concrete on two sides of the steel plate has the highest shear capacity and lowest drift. Concrete layers on both sides of the steel plate can completely prevent steel plate buckling. A numerical model by OpenSees was developed by Wang et al. [10] to study the cyclic behavior of SPRC composite shear walls. This work identified the effects of steel plate ratio, axial load ratio, concrete strength, and reinforcement ratio on the cyclic behavior of SPRC shear wall.
Although a number of experimental and numerical studies have been reported to investigate the general behavior of the SPRC shear wall, few of them illustrate the prediction of the deformation capacity of SPRC shear wall. Since displacementbased design has been a common approach to seismic design, the yield and ultimate displacement of the SPRC shear wall are critical indices to determine its performance levels. The main objective of this study is to develop an efficient nonlinear 3D finite element model to investigate the deformation of the SPRC shear wall under cyclic loads. ABAQUS was used in the analysis. Based on the finite element model which was validated against published test results, a parametric study of the SPRC shear wall with flexural failure was performed. Statistical analyses were utilized to propose formulas for yield rotation and ultimate rotation of SPRC shear walls.
2. Finite Element Modelling
Finite element software ABAQUS was used to simulate the behavior of SPRC shear walls. The main modelling approach is described as follows.
2.1. Modelling Approach
The SPRC composite shear wall studied in this paper is composed of four parts, concrete, structural steel, steel plate, longitudinal, and transverse reinforcement bars in wall web and boundary columns. The modelling of different parts of the SPRC shear wall is shown in Figure 2. Since structural steel and stirrups can provide different levels of confinement to concrete, the concrete part is divided into three zones accordingly, which are highly confined concrete, partially confined concrete, and unconfined concrete with different states of confinement. The highly confined concrete is taken from the web of the structural steel to the midwidth of each flange outstand. The unconfined concrete is taken from the centerline of the longitudinal reinforcement to the concrete cover. The partially confined concrete is the remaining part without any confinement, as shown in Figure 3 [11]. Their concrete model will be considered in different ways, which will be illustrated hereinafter.
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2.2. Finite Element Type and Modelling of Interfaces
The concrete was modelled by the 3D 8node solid elements with reduced integration C3D8R. The steel plate was modelled by using the 4node doubly curved shell elements with reduced integration S4R. The 2node linear displacement truss elements were used to model the steel rebars. Welldistributed studs welded to the steel plate can almost assure the coworking of concrete and steel plate [4]. Additionally, highstrength steel wires that pass through the reserved holes in steel plate is tied to the transverse rebars in concrete wall web to strengthen the connection between concrete and steel plate. As the tests show [7], these configurations can well control the concrete and steel plate to deform together. Thus, the possible sliding between concrete and steel plate is ignored. The steel plate, structural steel, and reinforcements are embedded into the concrete using the Embedded Constraint option in ABAQUS, assuming a perfect bond.
2.3. Material Modelling of Concrete
Concrete was modelled using the damaged plasticity model for concrete (CDP model) implemented in ABAQUS. The concrete strength and ductility under the confinement of steel and reinforcements are considerably improved. As illustrated by Ellobody and Young [12], the CDP model may not be able to accurately simulate the compressive behavior of concrete under a high level of confinement. In this case, a confined concrete model needs to be developed. The input concrete compressive strength f_{c} will be taken as the unconfined concrete cylinder compressive strength f_{c0}, while the corresponding concrete strain ε_{c} will be taken as the confined strain ε_{cc}. According to Yu et al. [13], parameters that control the yield surface and flow rule are altered for partially and highly confined concrete, respectively. For highly confined concrete, K_{c} = 0.725 and ψ = 56°. For partially confined concrete, K_{c} = 0.667 and ψ = 42°. To improve the convergence speed, a viscosity parameter μ equal to 0.001 is introduced.
2.3.1. Unconfined Concrete
The concrete compressive strain ε_{c} is taken as ε_{c0}, which is the unconfined concrete strain corresponding to the unconfined concrete strength f_{c0}, taken as 0.003 as recommended by the fib model code for concrete structures [14].
2.3.2. Partially Confined Concrete
The compressive strain ε_{c} of confined concrete can be determined by using equation (1), proposed by Mander et al. [15] and improved by Denavit et al. [16]:where f_{l} is the equivalent lateral confining pressure and f_{l1} and f_{l2} are the lateral confining pressures imposed by the reinforcement bars with different directions, respectively, as given by Mander et al. [15].
2.3.3. Highly Confined Concrete
High confinement of concrete is provided by the steel and reinforcements. The lateral confining pressure () imposed by the flange of the steel is given by Huang et al. [17]:where f_{ys} is the yield strength of the steel, L is the flange length, H is the web length, and t is the flange thickness of the steel. The confined strain ε_{c} of the highly confined concrete can be determined by the lateral confining pressure imposed by the reinforcements and steel in different directions.
The model proposed by Popovics [18] and Collins [19] was adopted to simulate the concrete behavior in compression. The compressive stress f_{c} and strain ε_{c} relation of the concrete is defined as follows:where f_{c} is the peak compressive strength of concrete, λ is the curvefitting factor, E_{c} is the initial tangent modulus, ε_{0} is the strain when σ_{c} reaches f_{c}, and d is the factor which controls the slope of the stressstrain curve.
The initial Young’s modulus of concrete is reasonably calculated using equation (13) given by the ACI Specification [20]. Poisson’s ratio of concrete is taken as 0.2:
Under uniaxial tension, the stressstrain response follows a linear elastic relationship until reaching the value of the failure stress. The tensile failure stress was assumed to be where f_{cu} is the cubic compressive strength of concrete. The softening stressstrain response, past the maximum tensile stress, was represented by a linear line defined by the fracture energy and crack bandwidth, as shown in Figure 4. The fracture energy G_{f} in N/mm (energy required to open a unit area of crack) was taken as equation (14) as recommended by the Fib Code [14]:
The concrete damaged plasticity model assumes that the reduction of the elastic modulus is given in terms of a scalar degradation variable d as in equation (15), where E_{0} is the initial modulus. Figure 5 shows the stressstrain relation for cyclic compressive loading:
The stiffness degradation variable d is a function of the stress state and the uniaxial damage variables d_{t} and d_{c}. According to Birtel and Mark [21], the damage variables d_{t} and d_{c} can be calculated by assuming a linear relation between plastic strain and inelastic strain:where b_{c} = 0.7 and b_{t} = 0.1, suggested by Birtel and Mark [21]. and are the compressive and tensile plastic strain of concrete, respectively.
2.4. Validation of Finite Element Model
To validate the proposed finite element model, numerical models of 9 SPRC shear walls with different shear span ratios were developed. The results of the numerical analyses were compared with that of the tests to illustrate the accuracy of the modelling method. The detailed parameters of the specimens are listed in Table 1.
 
h_{w}, l_{w}, and t_{w} are the height, length, and thickness of the composite shear wall; the concrete cylinder compressive strength f_{c} is taken as 0.76f_{cu} for C50, 0.82f_{cu} for C80 and f_{cu} multiplied by the linear interpolation between 0.76 and 0.82 for concrete with strength grade between C50 and C80; n is the axial load ratio; λ is the shear span ratio; f_{yt}, f_{yl}, f_{yp}, ρ_{t}, ρ_{l}, and ρ_{p} are the yield strengths and ratios of the transverse, longitudinal reinforcement, and steel plate of the wall web; f_{ys}, f_{ya}, ρ_{s}, and ρ_{a} are the yield strengths and ratios of the longitudinal reinforcement and structural steel of the boundary members. 
The loaddisplacement curves of the FEAs are compared with those of the tests in Figure 6. The softening of some specimens with a shear span ratio of 1.5 is not in good agreement with the test because the buckling of the steel plate at middle height is not well predicted in the numerical analyses. Even though the steel plate and concrete can almost deform together, the possible sliding between steel plate, reinforcement, and concrete will affect the unloading and reloading path of the cyclic behavior, as well as the idealization of the steel model. After all, the pinching effect, strength, and stiffness degradation are well predicted in most cases. The corresponding strengths and ductility ratios are listed in Tables 2 and 3. The equivalent energy method is adopted to obtain the yield point in the skeleton curve, and the ultimate displacement is taken as the displacement corresponding to the lateral force that decreases to 85% of the peak force. The strength and deformation capacities of the shear wall under cyclic loads are predicted with acceptable accuracy. It is concluded that well agreement between tests and numerical analyses is achieved.
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∆_{y} is the yield displacement of the specimen; ∆_{u} is the ultimate displacement of the specimen; μ is the ductility factor defined by the ratio of ultimate displacement to yield displacement. 
Two failure modes were observed in the test and confirmed in the finite element analysis as summarized in Table 2, that is, flexural failure (F) and flexuralshear failure (FS). Figures 7(a)–7(c) demonstrate the predicted and test concrete damage of specimen GB231 which shows a typical flexural failure. As shown in Figure 7(a), bottom concrete of boundary columns suffered the most severe compressive damage, which can be observed in the test as shown in Figure 7(c). Specimen W1a failed in a flexural and shear failure mode. Obvious shear cracks developed along the height of the shear wall, and the bottom concrete crushed under flexural and shear effects, as the test failure mode of specimen W1a shown in Figure 8(c). This can be confirmed in the FE results shown in Figures 8(a) and 8(b).
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3. Deformation Capacity of SPRC Shear Wall
Ductility is an important index of deformation capacity of members and structures in displacementbased design. Generally, ductility of members is represented by ductility factor defined as the ratio of ultimate displacement to yield displacement. Based on the test specimen in reference [7], 720 finite element models of the SPRC shear wall with potential flexural failure were developed. The typical section is shown in Figure 9. The main parameters of the FEMs are listed in Table 4.

The load history in most tests [5–7] begins with forcecontrolled elastic cycles to the predicted yield force, followed by displacementcontrolled cycles incremented by multiple of yield displacement. According to Rahnavard et al. [9], a simplified displacementcontrolled scheme is adopted to generate cyclic behavior of SPRC shear walls with different shear span ratios. The typical loading protocol is shown in Figure 10, which is represented by the drift angle defined as the lateral displacement divided by the total height of the specimen.
The yield strength and ultimate strength of steel and rebars are determined from the experiment, detailed values listed in Table 5. Cyclic analyses were conducted in ABAQUS. Then, the yield displacement and ultimate displacement were obtained in the same way above.

3.1. Yield Rotation of SPRC Shear Wall
Slender shear wall tends to be dominated by flexural failure. The yielding of a cantilever shear wall occurs when its bottom section yields, when the curvature is linearly distributed along the height of the wall. The curvature of the top section is 0. The yield curvature is taken as ϕ_{y}, and then the yield displacement of shear wall iswhere H is the height of the shear wall.
Priestley [22] suggests that the yield curvature of the shear wall is only relevant to the yield strain of tensile longitudinal reinforcement ε_{y} and the section depth h_{w}:
The yield rotation of the shear wall is given in the form where a is an undetermined coefficient:
By ignoring shear deformation, the yield rotation is defined as
The coefficient a is determined by regression analyses of finite element results. Then, the yield curvature and rotation are given by
The ratios of yield rotation calculated from finite element analyses and proposed formula are shown in Figure 8, with a mean value of 1.00 and coefficient of variation (C.V) of 33.4%. Results of shear walls with different steel plate ratios and axial load ratios are marked to illustrate the effect of these parameters. Experiment results suggest that the yield and ultimate deformations of the SPRC shear wall are improved compared with those of the steel reinforced concrete composite shear wall. The yield rotation tends to rise as the steel plate ratio increases, which can be seen in Figure 11(a). No obvious effect of axial load ratio is found on yield rotation of SPRC shear walls.
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3.2. Ultimate Rotation of SPRC Shear Wall
The ultimate rotation can be calculated on the basis of purely flexural behavior by the concept of plastic hinge and plastic hinge length L_{p} for the slender shear wall [14]. The plastic hinge assumes that the plastic curvature ϕ_{p} is lumped and uniformly distributed in the plastic hinge length, where ϕ_{p} = ϕ_{u} − ϕ_{y}. The ultimate rotation is calculated bywhere ϕ_{p} and ϕ_{u} are the yield curvature and ultimate plastic curvature of the bottom section, respectively.
Priestley and Paulay [23] suggest that the plastic hinge length L_{p} of the shear wall is related to its section depth and the height:where h_{w} is the section depth and H is the height of the shear wall.
It is assumed that the plastic hinge fails when the compressive zone of the concrete reaches the ultimate concrete strain. The Mander model, as simplified by Paulay and Priestley [23] regarding the calculation of the ultimate strain of confined concrete, was applied:where ε_{su} is the elongation at steel rupture, ρ_{s} is the volumetric ratio of confining steel, f_{yw} is the yield strength of confining steel, and f_{cc} is the compressive strength of confined concrete. According to Jiang, ε_{su} is taken as 0.12.
On the basis of the planesection hypothesis, the compressive section depth x_{n} and the corresponding relative compressive section depth ξ_{u} are determined by the force balance which is shown in Figure 12.
For the SPRC shear wall of the rectangular section, the force balance iswhere
Then, the relative compressive section depth of concrete iswhere A_{s}, , ρ, ρ′, f_{y}, and are the areas, reinforcement ratio, and yield strength of tensile and compressive rebar of the boundary members of the shear wall, respectively; A_{a}, , ρ_{a}, , f_{a}, and are the areas, reinforcement ratios, and yield strength of tensile and compressive structural steel of boundary members, respectively; ρ_{p} is the steel plate ratio, where ρ_{sp} = t_{sp}/b_{w}; ρ_{sw} is the longitudinal reinforcement ratio of the shear wall web; f_{p} and f_{yw} are the yield strengths of steel plate and rebars of the wall web. N is the axial load; and N_{sp} and N_{sw} are the axial force resisted by steel plate and reinforcement of the wall web, respectively.
The ultimate curvature is calculated by the ratio of the ultimate strain to the compressive section depth:
By substituting equations (23), (25), (26), and (31), the ultimate rotation is
The predicted ultimate rotation θ_{u,pred} calculated by using equation (32) is compared with the numerical ultimate rotation θ_{u,FEM} in Figure 13. The mean value of the ratio of numerical to predicted values θ_{u,FEM}/θ_{u,pred} is 1.08, with the coefficient of variation of 18.9%.
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As shown in Figure 9, θ_{u,FEM}/θ_{u,pred} is similarly distributed under different shear span ratios and concrete grades, while the ratio tends to degrade as the steel plate ratio and axial load ratio increase because the compressive section depth is affected by the steel plate ratio and axial load. For a slender shear wall with the planesection hypothesis, a larger axial load ratio will increase the compressive section depth of concrete and consequently decrease the ultimate curvature. Equation (32) tends to underestimate the compressive section depth of the shear wall with a small steel plate ratio and axial load ratio. To simplify the statistical procedure, equation (32) is transformed into a simple formula that includes the main variables, that is, compressive section depth, characteristic volumetric value of the boundary member, and shear span ratio:where C_{0}, C_{1}, C_{2}, and C_{3} are undetermined coefficients.
The numerical values of the ultimate rotation are used to determine the coefficients by regression analyses:
Numerical values of θ_{u} are compared with the predictions by equation (34) θ_{u,pred2} in Figure 14. The mean value of the ratios of numerical to predicted values θ_{u,FEM}/θ_{u,pred2} is 1.00, with the coefficient of variation of 9.9%. The maximum and minimum ratios are 1.2 and 0.64, respectively.
The deformation capacities predicted by using equations (23) and (34) are compared with test results [6–8] in Table 6. The mean value of the ratios of the test to predicted yield deformations ∆_{y,TEST}/∆_{y,pred} is 1.02, with the coefficient of variation of 8.3%. The mean value of the ratios of the test to predicted ultimate deformations ∆_{u,TEST}/∆_{u,pred} is 1.04, with the coefficient of variation of 11.5%. A good agreement is achieved between the test and predicted deformation capacities of SPRC shear walls.

4. Conclusions
This paper investigates the deformation capacity of SPRC shear walls under cyclic loads. A nonlinear 3D finite element model in ABAQUS was developed and validated against published experimental results. Then, a parametric study was conducted to obtain the yield rotation and ultimate rotation of SPRC shear walls with potential flexural failure. By statistical analyses, formulas for the yield rotation and ultimate rotation for SPRC shear walls are proposed. The cyclic behavior of SPRC shear walls can be predicted by using these formulas and related studies on shear capacities and hysteretic models [7–10], which can be used by engineers to conduct a displacementbased performance design of SPRC shear wall system.
The proposed formulas focused on SPRC shear walls with high shear span ratio and flexuredominated behavior in which shear studs are well distributed according to the Chinese code for design of composite structures. Further work will involve the SPRC shear walls with low shear span ratio and different configurations of shear connectors.
Data Availability
The data used to support the findings of this study are available from the corresponding author upon request.
Conflicts of Interest
The authors declare that there are no conflicts of interest regarding the publication of this paper.
Acknowledgments
The financial support provided by the National Key Research and Development Program of China (Grant no. 2017YFC1500701), the National Natural Science Foundation of China (Grant nos. 51638016 and 51608406), and the Fundamental Research Funds for the Central Universities (WUT: 2018IVB013) is gratefully acknowledged.
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Copyright
Copyright © 2019 Zhi Zhou et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.