Experimental and theoretical investigation into the shear behavior of interfaces in double-faced superposed shear walls

Abstract

The double-face superposed shear wall is a sandwich type of prefabricated concrete structures, which has two interfaces compared with traditional reinforced concrete shear walls. The shear performances of the interfaces is crucial for the integrity and seismic performance of the double-face superposed shear walls. However, little has been known regarding to the mechanical behavior of interfaces in the double-faced Superposed Shear Walls. This study investigates the shear behavior of double-face superposed shear wall through 27 double-face superposed shear specimens pullout tests, the effects of the configuration of interface connection reinforcement, interface connection reinforcement height and reinforcement ratio of interface connection reinforcement were analyzed. The strain distribution patterns of the interface connection reinforcement within the double composite surfaces were obtained through strategically placed strain gauges on the reinforcement. Based on the experimental results, a formula for the shear capacity of the interfaces in double-face superposed shear specimens has been developed.

Keywords

double-face superposed shear wall shear behavior experimental study interface connection reinforcement theoretical investigation

Introduction

The double-face superposed shear wall, developed from precast concrete sandwich panels, is a prefabricated concrete structure with advantages of fast construction speed, less on-site framework, low pollution, etc. (Gu et al., 2022; Xue et al., 2022). Recently, with the environmental problems being increasingly prominent, such shear wall structures have been widely used in high-rise residential and commercial buildings. The current research regarding double-face superposed shear wall mainly focuses on the seismic performance of shear walls (Gu et al., 2022; Jiang et al., 2021; Wang et al., 2022; Xue et al., 2022), the horizontal joints at connections between precast walls or between the precast wall and the foundation (Chong et al., 2019; Gu et al., 2020; Hu et al., 2023), and the performance of the connectors between the precast panels (Gu et al., 2022; He et al., 2023; Wang et al., 2022). It has shown that most of the damage in double-face superposed shear walls is either shear failure at the horizontal joint or bond failure at the interfaces between the precast layer and the cast-in-place layer (Gu et al., 2022; Wang et al., 2022; Xue et al., 2022) The double-face superposed shear walls mainly resist the in-plane lateral forces induced by wind load and earthquake actions, thus the shear performance of the interfaces is the key to ensuring the overall performance of the structures. The performance of reinforcement concrete interfaces has been experimentally examined by many researches (Choi et al., 1999; Christopher et al., 2017; Costa et al., 2018 ; Hofbeck et al., 1969; Mattock et al., 1975; Mattock and Hawkins 1972; Nagle and Kuchma 2007; Palieraki et al., 2021; Papanicolaou and Triantafillou 2002; Rahal et al., 2016; Seible and Latham 1990), and design equations were also proposed to calculate the shear resistance of the interfaces with various reinforcements (Harries et al., 2012; Mansur et al., 2008; Pruijssers 1988). Compared with traditional cast-in-place shear walls, there are two bonding interfaces between the precast layer and the cast-in-place layer in double-face superposed shear walls. Besides, connecting reinforcements such as steel truss are commonly used to connect the two precast concrete wythes. In this reason, the performance of interfaces in double-face superposed shear walls should be different form the traditional reinforcement concrete interfaces.

This paper is aimed to shed light on the shear performance of interfaces of the double-face superposed shear walls. A total of 27 double-face superposed specimens are examined through pushout tests. The failure mode of specimens with double interfaces is analyzed, the shearing performances of interface with different types of connecting reinforcement are discussed, and the effect of the ratio of the interface reinforcement on shear capacities is studied. Based on the test results, this paper proposes a theoretical formula to evaluate the shear capacity of double composite interfaces.

Test overview

Specimen design and construction

The typical solution of the push-out test was adopted in the experimental program. A total of 27 double-face superposed specimens were tested, which can be divided into 4 groups based on the research variables. Steel trusses are typically placed between the precast layers and the cast-in-place layer of double-face superposed shear walls as connecting reinforcements; therefore group 1 were specimens with steel trusses. To investigate the influence of the interface reinforcement ratio, the diameter of the web rebar varied from 6mm to 10 mm. The height of the truss was fixed to 150 mm in group 1 specimens. Group 2 specimens were designed to examine the influence of the truss height on the shear performance of the interfaces, and the truss heights were set to 130 mm and 170 mm for tests BH6-13 and BH6-17, respectively. To study the influence of different forms of connecting reinforcement, tests in group 3 were specimens with steel stirrups. The diameter of the web reinforcement bars ranged from 6 mm to 10 mm, and the corresponding reinforcement ratio was from 0.19% to 0.52%. For comparison purpose, Group 4 were specimens with no reinforcement. All test specimens were the same in dimensions. The total thickness was 200 mm, including two 50 mm precast layers and a 100 mm post-cast layer. The width and the height of the specimens were 200 mm and 400 mm, respectively. The height difference between the precast layers and the post-cast layer was 100 mm, leading to a 300mm-height interface. The diameter of the chord rebars were 12 mm for all reinforced test specimens. The details of the test specimens are provided in Table 1 and the configurations of each specimen are shown in Figures 1∼3.

Table 1.

Design parameters of specimens.

Group	Test label	Specimen quantity	Precast concrete strength grade	Intermediate concrete strength grade	Interface reinforcement form	Web rebar diameter (mm)	Reinforcement ratio (%)
Group 1	ZH-6	3	C40	C30	Truss	6	0.19
	ZH-8	3	C40	C30	Truss	8	0.36
	ZH-10	3	C40	C30	Truss	10	0.52
Group 2	BH6-13	3	C40	C30	Truss	6	0.19
Group 2	BH6-17	3	C40	C30	Truss	6	0.19
Group 3	ZG-6	3	C40	C30	Stirrup	6	0.19
	ZG-8	3	C40	C30	Stirrup	8	0.36
	ZG-10	3	C40	C30	Stirrup	10	0.52
Group 4	NJ	3	C40	C30	None	None	0

Figure 1.

Schematic diagram of specimens with truss and stirrup.

Figure 2.

Geometric dimensions and configurations of specimens with (a) truss and (b) stirrup.

Figure 3.

Variables of all specimens.

The double-face superposed specimens were produced in Baoye Prefabricated Component Factory. Firstly, the formwork was placed on the same platform as the double-face superposed shear wall manufactured, and the connecting reinforcements were arranged within the formwork. By inserting 15 mm, 25 mm and 35 mm radius buckles on the lower chord rebars, the specimens with different truss heights were constructed. The 1^st precast layer was then cast with the use of a concrete placing boom. The casting of the 2^nd precast layer was performed after 3 days curing in a curing room so that the 1^st stage concrete had enough stiffness. A 100 mm thick wooden block was placed on the formwork to ensure the thickness of the post-cast layer. The inner layer of the specimen was cast after 3 days in the curing room. The specimens were cured outdoors for 2 months before testing. Figure 4 shows the construction process of the double-face superposed test specimens.

Figure 4.

Manufacture procedures of specimens: (a) Placing the buckles and the steel trusses; (b) Pouring the first precast layer; (c) Pouring the second precast layer; (d) Pouring the inner layer concrete.

Material properties

The nominal strength grade of the precast layer and inner layer concrete was C40 and C30, respectively. For each layer concrete, nine concrete cubes were cast at the same time of each casting and were cured along with the test specimens. The concrete cubes were tested according to the standard for test methods of mechanical properties of ordinary concrete (GB/T 50,081-2002) (China Architecture and Building Press 2019) and the average results are listed in Table 2.

Table 2.

Mechanical properties of concrete (Unit: Mpa).

Concrete	1st layer	2nd layer	Inner layer
Cubic compressive strength f_cu	39.8	40.4	37.2

The interface connecting reinforcements were of HRB300 strength grade (yield strength of 400Mpa). The tensile test was conducted according to the Metallic Material Room Temperature Tensile Test Method (GB/T 228-2002) (Standards Press of China 2010) Three samples were prepared for each diameter reinforcement and the average yield strength, ultimate tensile strength, and modulus of elasticity are given in Table 3.

Table 3.

Mechanical properties of rebars (Unit: Mpa).

Steel rebars	Yield strength	Ultimate tensile strength	Modulus of elasticity
Φ6	497.9	614.5	2.83◊10⁵
Φ8	407.9	600.5	2.63◊10⁵
Φ10	443.9	596.9	1.90◊10⁵

Test setup and loading protocol

The test was completed in the State Key Laboratory of Civil Engineering Disaster Prevention of Tongji University. The test loading device is shown in Figure 5. A 200T servo actuator was used to apply the in-plane shear force. The loading rate was 0.01 mm/min. The test terminated when a load of less than 50% of the peak load was observed. Two displacement gauges were set near the loading end and the free end, one at each side. To deeply understand the role of the steel rebar in the shear performance of the interfaces, electrical strain gauges were arranged in the steel rebars. The layout of strain gauges is shown in Figure 6, where the first number of the strain gauge indicates the stage of casting, the letter indicates the type of the rebars (C-chord member, W-web member, S-stirrup member, V-vertical member), and the second number indicates the No. Of the strain gauges. The strain gauges in the chord/vertical rebars were embedded in the precast layer, and the web/stirrup strain gauges were placed at the interfaces between the precast and post-cast concrete.

Figure 5.

(a) Loading instruments; (b) Arrangement of LVDTs.

Figure 6.

Arrangement of strain gauges: (a) Truss form; (b)Stirrup form.

Test phenomenon

NJ specimens

The failure processes of non-reinforced specimens were similar. No obvious cracks were observed before failure. When the load increased to the ultimate load of 173.5 kN, shear failure occurred on the interface between the 1st precast layer and the inner layer, while the interface between 1st precast layer and the inner layer remained intact. The failure process of NJ specimens is shown in Figure 7. As can be seen, shear failure was occurred along the entire surface, and the failure process showed brittle failure characteristics.

Figure 7.

Failure process of NJ-1 specimen: (a) Initial loading; (b) Failure surface of precast layer; (c) Failure surface of cast-in-place layer.

ZH specimens

The failure processes of the ZH6, ZH8 and ZH10 specimens were similar. When the load increased to 100 kN, the first cracks appeared at the bottom of the 2-3 interface (the interface between the 2^nd precast layer and the inner layer), as shown in Figure 8(a). When the load reached 159 kN, the cracks extended to the top of the specimen, as shown in Figure 8(b). New cracks were observed on the 1-3 interface (the interface between the 1^st precast layer and the inner layer) as the load increased to about 217 kN, as shown in Figure 8(c). The cracks developed rapidly and extended to the top of the specimen, and the ultimate load of 308.9 kN was reached, as demonstrated in Figure 8(d). As the loading continued, the bearing capacity suddenly dropped to 216 kN. When the load was reduced to 180 kN, an apparent horizontal crack was observed in the second precast layer (i.e., the embedded side of the top horizontal rebar). As indicated in Figure 8(e), the crack width in 2-3 interface was significantly larger than that of the1-3 interface. The second precast layer was removed after the test terminated, and it was observed that the horizontal crack penetrated the entire section of the second precast layer, as shown in Figure 8(f).

Figure 8.

Failure process of ZH6-1 specimen: (a) Initial crack at 2-3 interface; (b) Crack penetration on 2-3 interface; (c) Initial crack at 1-3 interface; (d) Crack penetration on 1-3 interface; (e) Horizontal cracks on the 2nd precast layer; (f) Horizontal crack along the entire section.

ZG specimens

ZG specimens were constructed with steel stirrups as the connecting reinforcements. The failure processes of the ZG6, ZG8 and ZG10 specimens were similar. When the load increased to 146 kN, cracks first appeared at the bottom of the 2-3 interface, as shown in Figure 9(a). The cracks quickly penetrated to the top of the specimen after initiated, and the load reached 169 kN, as shown in Figure 9(b). When the load increased to about 258 kN, cracks appeared on the 1-3 interface, as shown in Figure 9(c). The cracks developed quickly to the top of the specimen, and the ultimate load of 349.5 kN was reached, as shown in Figure 9(d). The load gradually decreased as the loading continued. The crack width on the 2-3 interface was similar to that on the 1-3 interface, as shown in Figure 9(e). Oblique cracks could be observed at the precast layers by the end of the test, as shown in Figure 9(f).

Figure 9.

Failure process of ZG6-1 specimen: (a) Initial crack at 2-3 interface; (b) Crack penetration on 2-3 interface; (c) Initial crack at 1-3 interface; (d) Crack penetration on 1-3 interface; (e) Comparison of interface crack forms.

BH specimens

The failure process of the BH specimens was similar to that of the ZH specimens. Due to insufficient anchorage of the BH6-13-3 specimen, the shear bearing capacity could only reach half of the bearing capacity of BH6-13-1 and BH6-13-2. The damage was mainly due to insufficient anchoring depth of the upper chord reinforcement. The crack on the 2-3 interfaces was obviously larger than that of 1-3 interfaces, as shown in Figure 10(a). After the test, the concrete specimen was peeled off. As demonstrated in Figure 10(b), there was very little concrete wrapped around the chord rebars on the 2-3 interface.

Figure 10.

Failure process of BH6-13-1 specimen: (a) Cracks at peak; (b) The chord rebar pulling out.

In summary, all the double-face superposed specimens suffered shear failure at the precast and post-cast interface. Shear failure of the non-reinforced specimens was occurred along the entire surface, and the failure process showed brittle failure characteristics. While duet to the existence of the connecting reinforcements, shear failure of the reinforced specimens was much more moderate. The specimens reached their peak point when cracks appeared on both interfaces and penetrated to the top of the specimens. Table 4 summarizes the order of the cracks appeared. As can be seen, the first crack appeared randomly on the 1-2 interface or the 2-3nterface. The reason for this is that the bonding force of the interface was affected by many factors, such as the interface roughness, the concrete vibration quality during pouring, and the cubic compressive strength of new and old concrete, etc.

Table 4.

Failure modes of specimens.

Specimen label	First crack (cracking load kN)	Rear crack (cracking load kN)	Peak load (kN)
NJ-1	1-3 interface (173.5)	-	173.5
NJ-2	1-3 interface (171.3)	-	171.3
NJ-3	2-3 interface (121.5)	-	121.5
BH6-13-1	1-3 interface (154.7)	2-3 interface (213.5)	314.7
BH6-13-2	2-3 interface (150.6)	1-3 interface (228.7)	313.2
BH6-13-3	1-3 interface (132.5)	2-3 interface (158.5)	158.5
BH6-17-1	1-3 interface (164.5)	2-3 interface (247.3)	329.8
BH6-17-2	1-3 interface (131.6)	2-3 interface (201.3)	315.4
BH6-17-3	1-3 interface (171.3)	2-3 interface (171.3)	320.7
ZH6-1	1-3 interface (100.3)	2-3 interface (217.6)	308.9
ZH6-2	2-3 interface (120.4)	1-3 interface (197.6)	290.8
ZH6-3	2-3 interface (206.3)	1-3 interface (295.6)	351.9
ZH8-1	1-3 interface (177.3)	2-3 interface (247.6)	308.4
ZH8-2	1-3 interface (157.3)	2-3 interface (261.6)	296.4
ZH8-3	1-3 interface (203.1)	2-3 interface (301.2)	386.1
ZH10-1	1-3 interface (270.1)	2-3 interface (303.4)	315.1
ZH10-2	1-3 interface (290.3)	2-3 interface (308.5)	314.6
ZH10-3	1-3 interface (268.9)	2-3 interface (347.2)	365.1
ZG6-1	2-3 interface (146.7)	1-3 interface (245.6)	347.1
ZG6-2	2-3 interface (106.4)	1-3 interface (195.7)	310.8
ZG6-3	1-3 interface (126.3)	2-3 interface (186.6)	305.4
ZG8-1	2-3 interface (176.1)	1-3 interface (266.2)	349.9
ZG8-2	2-3 interface (157.4)	1-3 interface (243.2)	331.5
ZG8-3	2-3 interface (161.5)	1-3 interface (252.4)	341.7
ZG10-1	2-3 interface (257.1)	1-3 interface (308.2)	375.9
ZG10-2	2-3 interface (274.4)	1-3 interface (301.4)	365.7
ZG10-3	2-3 interface (251.2)	1-3 interface (298.2)	357.3

Analysis of test results and discussions

Comparisons of reinforced specimens and non-reinforced specimens

It can be seen from the failure processes in the previous section that the failure of the double-face superposed non-reinforced specimen exhibited obvious brittle failure characteristics. When the crack was observed, the load reached the peak value, and the interface was completely sheared. The failure of the double-face superposed specimen with reinforcement was a ductile failure, and cracks on both interfaces could be observed. The peak loads of all specimens are provided in Table 4. As can be seen, the shear capacities of ZH6 and ZG6 specimens were almost twice of that of the NJ specimens. The load-displacement curves of the free ends of specimens NJ-1, ZH6-1 and ZG6-1 are shown in Figure 11, from which we can see that the brittle failure of the double-face superposed non-reinforced specimens was obvious, and the ductility of double-face superposed specimens with reinforcement was better than the non-reinforced specimens.

Figure 11.

Load-displacement curves of the reinforced specimen and non-reinforced specimen.

The influence of truss height

The failure modes of double-face superposed specimens with heights ranging from 130 mm to 170 mm was basically the same. The load-displacement curves of specimens BH6-13-1, ZH6-1 and BH6-17-1 are shown in Figure 10. It can be seen from the figure that the change in the height of truss had no obvious influence on the load-displacement curves. Regardless of the BH6-13-3 specimens due to insufficient anchorage, the peak loads of the BH6-13, BH6-17 and ZH6 specimens are shown in Table 5. As shown in Table 5, the effect of the height of truss rebars on the shear capacity of interfaces was small. (Figure 12).

Table 5.

The ultimate capacity of specimen with different truss heights (kN).

Specimen number	BH6-13	ZH6	BH6-17
Specimen1	314.7	308.9	329.8
Specimen2	313.2	290.8	315.4
Specimen3	-	351.9	320.7
Average value	313.9	317.2	322.0

Figure 12.

Comparisons of load-displacement curves of specimens with different truss heights.

The influence of different forms of connecting reinforcement

The failure process of the double-face superposed specimens of the truss-type connecting reinforcement and the stirrup-type connecting reinforcement were basically the same before the peak load. The cracks were first initiated on one of the interfaces, and then developed rapidly and penetrated through the interface. As the load increased, new cracks appeared on the other interface. The cracks then extended through the interfaces until the load reached its peak. The difference between the two types was observed after the peak load: (1) As the displacement continued to increase, the widths of the cracks on the two interfaces were different for the truss type connecting reinforcement. The cracks on the 2-3 interface were generally larger than those on 1-3 interfaces, no matter which side was the cracks initiated on. Horizontal cracks were observed on the outer side of the second precast surface. The width of the cracks on both sides of the specimen with the stirrup type connecting reinforcement was not much different. No horizontal crack occurred on the precast layers. (2) After the peak load, the load of the specimens with truss type connecting reinforcement suddenly dropped. While the load of the specimens with stirrup type connecting reinforcement decreased slowly, and the reduction was relatively small. The load-displacement curves at the free ends of specimens ZH6-1 and ZG6-1, ZH8-1, ZG8-1, ZH10-1, and ZG10-1 are shown in Figure 13. As can be seen, before the peak load was reached, the slipping displacement at the free end of specimens with two types of connecting reinforcement did not exceed 2 mm. After the peak load was reached, the load of the specimen of connecting reinforcement in the truss form was reduced sharply. Meanwhile, the increment of slipping displacement was small. The load of the specimens of connecting reinforcement in the form of stirrup gradually decreased with the increase of the slipping displacement, which showed good ductility.

Figure 13.

Load-Displacement curves of specimens with different forms of connecting reinforcements: (a) ZH6-1 specimen and ZG6-1 specimen; (b) ZH8-1specimen and ZG8-1 specimen; (c) ZH10-1specimen and ZG10-1 specimen.

The influence of interface steel reinforcement ratios

As indicated in the damaging process, when the interface reinforcement ratio increased from 0.19% to 0.52%, the cracking load of the specimens (independent of connecting reinforcement type) increased with the reinforcement ratio. However, the failure modes did not change substantially. The load-displacement curves of specimens ZH6-1, ZH8-1 and ZH10-1 and specimens ZG6-1, ZG8-1 and ZG10-1 are shown in Figure 14. The effects of the reinforcement ratio of connecting reinforcements were not obvious, as presented in the graphs. The peak loads of specimens ZH6, ZH8, and ZH10 are shown in Table 6. The peak loads of specimens ZG6, ZG8, and ZG10 are shown in Table 7. When the interface reinforcement ratio increased from 0.19% to 0.52%, the load-bearing capacity of the specimens with truss type connecting reinforcement and stirrup type connecting reinforcement increased by 4.5% and 14.1%, respectively.

Figure 14.

Load-displacement curves of specimens with different interface steel reinforcement ratios: (a) Specimens with truss connecting reinforcements; (b) Specimens with stirrup connecting reinforcements.

Table 6.

The ultimate capacity of specimens with truss connecting reinforcements (kN).

Specimen number	ZH6	ZH8	ZH10
Specimen1	308.9	296.4	315.1
Specimen2	290.8	308.4	314.6
Specimen3	351.9	386.1	365.1
Average value	317.2	330.3	331.6

Table 7.

The ultimate capacity of specimens with stirrup connecting reinforcements (kN).

Specimen number	ZG6	ZG8	ZG10
Specimen1	347.1	349.9	375.9
Specimen2	310.8	331.5	365.7
Specimen3	305.4	341.7	357.3
Average value	321.1	341.0	366.3

Strain development in connecting reinforcements

ZH6 specimens

The strain gauges arranged on the connecting reinforcement of ZH6-1 measured the strain of the steel rebars with the load, as shown in Figure 15. The strain developments of steel rebars on the same interface before peak load were relatively consistent. The strain of the steel rebar on the first cracking side developed quicker, as opposed to that of the later cracking side. On the 1-3 interface where cracks first initiated, the rebar strain was about 3× $10^{- 4}$ before the load of 100 kN was reached. The cracks passed through the layer at about 150 kN. At this point, the strain of steel rebar increased rapidly and exceeded the yield strain. The strain of steel rebar on the uncracked side (2-3 interface) was still small (about 5 × 10⁻⁴). When the load was increased to 220 kN, cracks were observed on the 2-3 interface, and the strain of steel rebar was about 1.5 × 10⁻³. As the load increased, the strain of the steel rebar at the interface increased slowly. At the peak load, the strain of the steel rebar on this surface was about 1.6 × 10⁻³, which has not reached the yield strain. The bearing capacity of the specimen dropped suddenly after the peak load. At this stage, the strain of the web rebar increased sharply and reached 6.6 × 10⁻³. The strains of the chord rebars were small during the entire loading process, and the maximum strain did not exceed 2 × 10⁻³.

Figure 15.

Strain distribution of steel rebar of ZH6-1: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces- the loading end; (d) Strain distribution of different interfaces- the free end.

ZG6 specimens

The relationship between the horizontal force and the strain of the connecting reinforcements of ZG6-2 specimen is shown in Figure 16. The strain developments of steel rebars on the same interface before peak load were relatively consistent. The strain of steel rebars on the first cracking side reached the yield strain before peak load, and so did those on the later cracking side. On the 2-3 interface where the first crack was initiated, the strain of rebar was about 7 × 10⁻⁴ before a load of 90 kN was reached. Then, the strain of rebars on the 2-3 interface rapidly increased beyond the yield strain at about 100 kN. Meanwhile, the strain of rebar on the 1-3 interface (i.e., uncracked side) was small with a value of about 3.5 × 10⁻⁴. Until the load increased to 190 kN, new cracks were observed on the 1-3 interface, and the strain of the rebar was about 6 × 10⁻⁴. As the load increased, the crack penetrated through the 1-3 interface. The strain of the steel on this side increased significantly to 4.5 × 10⁻³, which exceeded the yield strain. At peak load, the strain corresponding to 1S3 strain gauge was about 1.7 × 10⁻². The strain values of the vertical members during the whole loading process were small, with a maximum strain of below 5 × 10⁻⁴.

Figure 16.

Strain distribution of steel rebar of ZG6-2: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces- the loading end.

ZH8 specimens

The relationship between the horizontal force and the strain of the connecting reinforcements of ZH8-2 specimen is shown in Figure 17. The strains of the steel rebars of ZH8-2 had the same pattern as the those of ZH6-1. Before the peak load, the strain developments of steel rebars on the same interface were relatively consistent. The strain of steel rebars on the first cracking side developed faster compared to that on the later cracking side. Further, the strain of steel rebar on the first cracking side reached the yield strain whereas that on the later cracking side did not.

Figure 17.

Strain distribution of steel rebar of ZH8-2: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces- the loading end.

ZG8 specimens

Figure 18 shows the relationship between the horizontal force and the strain of the connecting reinforcements of ZH8-2 specimen. The strain developments of the steel rebars of ZG8-1 had the same pattern as those of ZG6-1. It can be seen from the figure that before the peak load, the strain developments of the steel rebars on the same interface were relatively consistent. The strain of steel rebar near the loading end reached the yield strain before the peak load.

Figure 18.

Strain distribution of steel rebar of ZG8-1: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces-the loading end.

ZH10 specimens

The relationship between the horizontal force and the strain of the connecting reinforcements of ZH10-2 specimen is shown in Figure 19. The strain of steel rebar of the ZH10-2 specimen had the same development as ZH6-1 and ZH8-2. Before the peak load, the strain developments of the steel rebars on the same interfaces were relatively consistent. The strain of the steel rebar on the first cracking side grew quicker than that on the later cracking side. Overall, the strain values of steel rebars at the interface did not exceed the yield strain. The strain of steel rebar near the loading end on the first cracking side was close to the yield strain.

Figure 19.

Strain distribution of steel rebar of ZH10-2: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces-the loading end; (d) Strain distribution of different interfaces-the free end.

ZG10 specimens

Figure 20 shows the relationship between the horizontal force and the strain of the connecting reinforcements of ZG10-2 specimen. The strain of the steel rebar of the ZG10-2 exhibited the same trend as the test pieces ZG6-2 and ZG8-1. Before the peak load, the strain developments of the steel rebars on the same interface were relatively consistent. The strain of the steel rebar near the loading end reached the yield strain before the peak load while and the strain of the steel rebar near the free end did not reach the yield strain.

Figure 20.

Strain distribution of steel rebar of ZG10-2: (a) Strain distribution of the rear cracking side; (b) Strain distribution of the first cracking side; (c) Strain distribution of different interfaces - the loading end; (d) Strain distribution of different interfaces - the free end.

Analysis of shear bearing capacity of the bonding surface of the double-faced superposed shear specimen

According to the theory of shear friction, the shear bearing capacity of the composite interface is primarily determined by three key factors: interfacial bond strength, interfacial friction, and the dowel effect of interface reinforcement. Through a comprehensive analysis of these components, this study proposes a calculation formula for assessing the shear bearing capacity of the composite interface in double-sided composite specimens.

Interfacial bond strength V_adh (s)

The interfacial bonding force $V_{a d h} (s)$ is primarily influenced by two critical factors: the bond strength between new and existing concrete, and the bonded area of the interface. In accordance with the design of concrete structures of Eurocode 2 and the 1990 CEB-FIP Model Code , the bond strength can be assumed to be equivalent to the tensile strength of the concrete $f_{t}$ . The experimental results indicate that, in the ultimate state, one of the bonding surfaces of the double-sided composite specimens has cracked, suggesting that both bonding surfaces do not act fully simultaneously. Assuming an effective area coefficient $ς$ for the double-sided composite specimens in the ultimate state, the interfacial bonding force can be calculated using formula (1), where $A$ represents the total area of the entire bonding surface.

V_{a d h} (s) = 2 ς A f_{t}

(1)

Upon completion of the test, the ZH8 and ZG8 specimens were subjected to a stripping process. The interface conditions are depicted in Figure 21. As illustrated in the figure, after the load on the specimens decreased to 50% of the peak load, the concrete within the composite interface, confined by the truss reinforcement/stirrups, exhibited an excellent bonding condition. Based on the test phenomena, it can be assumed that at peak load, the effective bonding zone of ZG specimens encompasses the stirrup limb spacing and extends one concrete cover thickness outward from the outer surface of the stirrup reinforcement. For the ZH specimens, owing to the asymmetry in the connecting reinforcement, the effective area on the side of the upper chord reinforcement is defined as the region encompassing 0.5 times the limb spacing of the lower chord reinforcement and extending one concrete cover thickness from the outer surface of the reinforcement,as shown in Figure 22. The effective area coefficient $ς$ can be determined using Equations (2) and (3), where h represents the height of the superimposed surface.

F o r Z H s p e c i m e n s ς = (1.5 c_{1} + 4 c) / 2 h

(2)

F o r Z G s p e c i m e n s ς = (c_{1} + 2 c) / h

(3)

Figure 21.

The interface condition of the specimen after delamination. (a) The 2-3 interface of ZH8 specimen. (b) The 1-3 interface of ZH8 specimen. (c) The 2-3 interface of ZG8 specimen (d) The 1-3 interface of ZG8 specimen.

Figure 22.

The effective bonding area.

Interfacial friction V_sf (s)

The interfacial frictional force $V_{s f} (s)$ at the composite surface is derived from the tensile action of the interface connecting reinforcement. The frictional force can be calculated using formula (4). $F_{s}$ denotes the tensile force exerted on the interface connecting reinforcement perpendicular to the composite surface, while $μ$ represents the friction coefficient of the composite surface. Due to the rough treatment of the interface of the composite specimens, the friction coefficient $μ$ is taken as 1 according to the design of concrete structures of Eurocode 2 and the CEB-FIP design code of 1990.

V_{s f} (s) = μ F_{s}

(4)

The test results indicate that the anchorage performance of the interface reinforcement remains satisfactory upon composite surface failure. Data obtained from strain gauges arranged on the surface of the reinforcement show that, under ultimate load conditions, as the reinforcement diameter increases, the stress in the interface connection reinforcement does not fully reach yield strength. Consequently, in the ultimate state, the tensile force of the interface connection reinforcement $F_{s}$ can be determined by calculating the stress in each individual interface reinforcement, as presented in equation (5).

F_{s} = \sum_{i = 1}^{n} f_{s i} A_{s}

(5)

f_{s i}

represents the stress in each individual interface reinforcement at the peak load,

n

represents the number of reinforcing bars passing through the composite surface, and

A_{s}

represents the cross-sectional area of the interface connection reinforcement. While, the truss-form reinforcing bars are not perpendicular to the composite surface, it is necessary to convert the tensile force of the truss reinforcing bars

F_{s y c}

. The force condition is illustrated in Figure 23. Based on the spatial angle of the truss reinforcing bars,

F_{s y c}

is transformed into the tensile force

F_{s}

perpendicular to the composite surface, as presented in equation (6).

α

and

θ

respectively represent the spatial angles of the truss reinforcement, and their calculation methods are shown in formulas (7) and (8),

h_{c}

represents the truss height,

L

represents the length of the web bar, and

c_{1}

and

c_{2}

respectively represent the limb spacing and limb width of the truss reinforcement.

F_{s} = \sum_{i = 1}^{n} f_{s i} A_{s} \cos α \cos θ

(6)

\cos α = \frac{\sqrt{L^{2} - {(c_{2} / 2)}^{2}}}{L}

(7)

\cos θ = \frac{h_{c}}{\sqrt{{(c_{1} / 2)}^{2} + {(h_{c})}^{2}}}

(8)

Figure 23.

Schematic diagram of force distribution on truss reinforcement.

Let $A_{s e} = A_{s} \cos α \cos θ$ . Consequently, equation (6) can be rewritten as presented in equation (9). For the interface connection reinforcement in the form of stirrups, $\cos α = \cos θ = 1$ .

F_{s} = \sum_{i = 1}^{n} f_{s i} A_{s e}

(9)

At peak load, the interface reinforcement stresses calculated from strain gauge measurements on the steel bar surfaces are presented in Tables 8 and 9. The last column in the tables represents the average stress of the interface reinforcement. The data indicate that the average interface stresses in both ZH specimens and ZG specimens demonstrate a consistent variation pattern corresponding to increases in reinforcement diameter.

Table 8.

The stress of the interface connecting reinforcement of ZH specimens at peak load (MPa).

Test label	The stress from strain gauge 1w1	The stress from strain gauge 1w3	The stress from strain gauge 2w2	The stress from strain gauge 2w4	The average stress
ZH6-1	407.2 (0.82 $f_{y}$ )	468.3 (0.94 $f_{y}$ )	497.9 ( $f_{y}$ )	497.9 ( $f_{y}$ )	467.8 (0.94 $f_{y}$ )
ZH8-2	407.9 ( $f_{y}$ )	407.9 ( $f_{y}$ )	311.4 (0.76 $f_{y}$ )	264.3 (0.64 $f_{y}$ )	347.9 (0.85 $f_{y}$ )
ZH10-2	386.9 (0.87 $f_{y}$ )	219.6 (0.49 $f_{y}$ )	214.0 (0.46 $f_{y}$ )	205.4 (0.48 $f_{y}$ )	256.5 (0.58 $f_{y}$ )

Table 9.

The stress of the interface connecting reinforcement of ZH specimens at peak load (MPa).

Test label	The stress from strain gauge 1S1	The stress from strain gauge 1S3	The stress from strain gauge 2S2	The stress from strain gauge 2S4	The average stress
ZG6-1	497.9 ( $f_{y}$ )	497.9 ( $f_{y}$ )	497.9 ( $f_{y}$ )	497.9 ( $f_{y}$ )	497.9 ( $f_{y}$ )
ZG8-1	407.9 ( $f_{y}$ )	265.2 (0.65 $f_{y}$ )	389.5 (0.95 $f_{y}$ )	-	354.2 (0.87 $f_{y}$ )
ZG10-2	410.5 (0.92 $f_{y}$ )	178.3 (0.40 $f_{y}$ )	395.9 (0.89 $f_{y}$ )	39.3 (0.08 $f_{y}$ )	256.0 (0.58 $f_{y}$ )

Given that the average stress of the interface reinforcement correlates with variations in reinforcement diameter, multiplying both the average stress $f_{s}$ and yield strength $f_{y}$ by the reinforcement area $A_{s e}$ enables derivation of the variation trends for the average tensile force ( $f_{s} A_{s e}$ ) and yield tensile force ( $f_{y} A_{s e}$ ) in the interface reinforcement of ZH and ZG specimens, as depicted in Figure 24. The figure demonstrates that the variations exhibit nonlinear behavior. By applying a quadratic function to fit the experimental data, the relationships between $f_{s} A_{s e}$ and $f_{y} A_{s e}$ in ZH specimens and ZG specimens can be determined, with the corresponding mathematical formulations provided in Equations (10) and (11).

Z H s p e c i m e n s f_{s} A_{s e} = - 0.02 {(f_{y} A_{s e})}^{2} + 1.2 f_{y} A_{s e} (k N)

(10)

Z G s p e c i m e n s f_{s} A_{s e} = - 0.02 {(f_{y} A_{s e})}^{2} + 1.3 f_{y} A_{s e} (k N)

(11)

Figure 24.

The trend chart of $f_{s} A_{s e}$ and $f_{y} A_{s e}$ Dowel effect of interface reinforcement $V_{s r} (s)$ .

The correlation coefficients R of Formulas (10) and (11) are 0.9981 and 0.9902 respectively, indicating that the regression equations are significant. By dividing both sides of Equations (10) and (11) by $A_{s e}$ , the calculation formulas for the average stress $f_{s}$ of the interface connecting reinforcement of the ZH specimen and the ZG specimen can be derived, as shown in Equations (12) and (13), respectively.

Z H s p e c i m e n s f_{s} = - 2.0 \times 10^{- 5} A_{s e} {(f_{y})}^{2} + 1.2 f_{y} (M p a)

(12)

Z G s p e c i m e n s f_{s} = - 2.0 \times 10^{- 5} A_{s e} {(f_{y})}^{2} + 1.3 f_{y} (M p a)

(13)

Shear stresses induced in the interface connecting reinforcement due to interfacial slip generate dowel action. Based on the analytical framework established by Rasmussen (Rasmussen B H 1963) and Dulacska (Dulacska, 1972), while incorporating the angular orientation of truss reinforcement and rebar stress parameters, the dowel action force $V_{s r} (s)$ at the composite interface can be mathematically described by equation (14).

V_{s r} = \sum_{i = 1}^{n} 1.3 d^{2} \sqrt{{f_{c}}^{'} f_{s i}} \cos α \cos θ

(14)

{f_{c}}^{'}

represents the compressive strength of concrete cylinders, and

d

represents the diameter of the interface reinforcement.

Based on the aforementioned analysis, the shear bearing capacity of the composite interface in double-sided composite specimens can be expressed by equation (15).

V = 2 ς f_{t} A + n f_{s} A_{s e} + 1.3 n d^{2} \sqrt{{f_{c}}^{'} f_{s}} \cos α \cos θ

(15)

The proposed formula was employed to calculate the shear capacity of the ZH and ZG specimens, and the calculated values were compared with the minimum values from each group of test results. The results are shown in Table 10. As can be observed from the table, the calculated results using the proposed formula exhibit good agreement with the experimental data. For the ZH6 and ZG6 specimens, there is a significant discrepancy between the calculated values and the experimental results. This discrepancy arises because, for specimens with smaller interface reinforcement diameters, the interface shear capacity is primarily determined by the bond strength at the composite surface. However, the formula employs a relatively conservative approach in defining the effective area coefficient, leading to an underestimation of the calculated values.

Table 10.

The comparison of calculated values and experimental values.

Test label	The experimental value (kN)	The calculated value (kN)	The calculated value/The experimental value
ZH6-1	290.8	195.4	0.67
ZH8-2	296.4	244.6	0.83
ZH10-2	314.6	282.5	0.90
ZG6-1	305.4	236.9	0.76
ZG8-1	331.5	289.7	0.87
ZG10-2	357.3	330.8	0.92

Conclusions

To investigate the shear performance of interfaces between precast layers and cast-in-place layers in double-face superposed shear walls, this study conducted push-out tests on 27 double-face superposed specimens featuring various types of connecting reinforcement. The failure modes and shear capacities of these specimens were analyzed and discussed in detail. Based on the shear friction theory, this paper develops a formula for calculating the shear bearing capacity of the interface in double-sided composite specimens. The primary conclusions are summarized as follows:

1. All double-face superposed specimens showed shear failures on the interfaces. The failure process of the double-face superposed specimens without reinforcement exhibited obvious brittle failure characteristics. No obvious cracks were observed before failure. One of the two interfaces was sheared when damaged, while the other remained intact. Shear failure of the reinforced specimens was moderate duet to the existence of the connecting reinforcements. The occurrence of cracks at the two interfaces of the reinforced specimen was not synchronized. The cracks were initiated on one of the interfaces, and followed by those on the other as the load increased. The specimens reached their peak when cracks appeared on both interfaces and penetrated to the top of the specimens.

2. There was no obvious difference between the failure process of the double-face superposed specimens with truss rebars and stirrups before the peak load. While after the peak load, the bearing capacity of the double-face superposed specimens with the truss rebars dropped suddenly. The crack width of the 2-3 interface was obviously larger than that of the 1-3 interface. After the double-face superposed specimen in the form of stirrups reached the peak load, its bearing capacity decreased slowly with the increase of displacement, and the crack width on both sides of the interface was similar.

3. On the premise that the connecting reinforcement is well anchored, the shear bearing capacity of the interfaces of the double- superimposed specimen is less affected by the change in height of the truss reinforcement.

4. When the interface reinforcement ratio increased from 0.19% to 0.52%, the cracking loads of the double-face superposed specimens with two types of connecting reinforcement (truss-type and stirrup type) increased with the reinforcement ratio. However, the failure mode did not change substantially. When the interface reinforcement ratio increased from 0.19% to 0.52%, the shear capacity of the double-face superposed specimens of the truss-type connecting reinforcement increased by 4.5%, and the shear capacity of the double-face superposed specimens of the stirrup-type connecting reinforcement increased by 14.1%.

5. The strain development of the web members on the same interface were quite consistent. The connecting reinforcement in the form of stirrups was more effective than the connecting reinforcement in the form of trusses. For the double-face superposed specimens connected by the truss, the strain of rebar on the first cracking side developed quicker than that of the rear cracking side. The strain of the steel rebar on the rear cracking side did not reach the yield strain even at the peak load. For the double-face superposed specimens with stirrup connections, the strain of steel rebar near the loading end reached the yield strain before the peak load.

6. Based on the shear friction theory, this paper proposes a formula for calculating the shear bearing capacity of the bonding surface of double-sided composite specimens. Additionally, this study derives calculation formulas for the effective area coefficient and the average stress of interface reinforcement under limit states, which are then applied to the calculation of the shear bearing capacity of the interface. The research results show that the the proposed theoretical formulas exhibit excellent agreement with experimental results.

Footnotes

ORCID iD

Wenying Zhang

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: This research was supported by China Postdoctoral Science Foundation (Grant No.2022T150612). Any opinions, findings, conclusions or recommendations expressed in this article are those of the authors and do not necessarily reflect the views of the sponsors.

Declaration of conflicting interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Data Availability Statement

Some or all data used during the study are available from the corresponding author upon request. The data include experimental data, and experimental and computational results.*

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Experimental and theoretical investigation into the shear behavior of interfaces in double-faced superposed shear walls

Abstract

Keywords

Introduction

Test overview

Specimen design and construction

Material properties

Test setup and loading protocol

Test phenomenon

NJ specimens

ZH specimens

ZG specimens

BH specimens

Analysis of test results and discussions

Comparisons of reinforced specimens and non-reinforced specimens

The influence of truss height

The influence of different forms of connecting reinforcement

The influence of interface steel reinforcement ratios

Strain development in connecting reinforcements

ZH6 specimens

ZG6 specimens

ZH8 specimens

ZG8 specimens

ZH10 specimens

ZG10 specimens

Analysis of shear bearing capacity of the bonding surface of the double-faced superposed shear specimen

Interfacial bond strength Vadh (s)

Interfacial friction Vsf (s)

Conclusions

Footnotes

ORCID iD

Funding

Declaration of conflicting interests

Data Availability Statement

References

Interfacial bond strength V_adh (s)

Interfacial friction V_sf (s)