push-out tests for shear connectors in gfrp-concrete

Push-Out Tests for Shear Connectors in GFRP-Concrete Composite Bridge Deck SlabsHailin Huang, Ao Li, Lin Chen, Chuijun Zeng and Mingqiao ZhuJournal of Advanced Concrete Technology, volume 16 (2018), pp. 368-381

Predicting Bond Strength between FRP Plates and Concrete Sub-strate: Applications of GMDH and MNLR ApproachesSeyed Mahmood Hamze-Ziabari, Amir YasavoliJournal of Advanced Concrete Technology, volume 15 (2017), pp. 644-661

Study on Load-Slip Characteristic Curves of Perfobond Shear Connectors in Hybrid StructuresWei'an Wang, Canhui Zhao, Qiao Li, Weilin ZhuangJournal of Advanced Concrete Technology, volume 12 (2015), pp. 413-424

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Journal of Advanced Concrete Technology Vol. 16, 368-381, August 2018 / Copyright © 2018 Japan Concrete Institute 368

Scientific paper

Push-Out Tests for Shear Connectors in GFRP-Concrete Composite Bridge Deck Slabs Hailin Huang1*, Ao Li2, Lin Chen3, Chuijun Zeng3 and Mingqiao Zhu4

Received 12 January 2018, accepted 15 August 2018 doi:10.3151/jact.16.368

Abstract This study investigates structural behavior of shear connectors in glass fiber-reinforced polymer (GFRP)-concrete composite bridge deck slabs. Five kinds of push-out specimens with GFRP shear connectors were prepared and tested. The failure modes of the specimens were observed, and the load-slip curves were obtained. The effects of rib shapes, presence of holes, and existence of transverse rebars on failure mode, shear resistance, and ductility were investigated. Results showed that (1) the shear resistance and ductility of T-type rib shear connectors were greater than those of flat plate shear connectors. (2) Shear resistance and ductility were significantly improved by holes. (3) The transverse rebar has a significant influence on the failure mode. The concrete slabs in the specimens without a transverse rebar failed. Similarly, the GFRP shear connector in the specimens with a transverse rebar collapsed. (4) All of the specimens showed a similar failure mode caused by the shearing failure of the connectors in the root when the transverse rebar was applied. Furthermore, the shear resistance and ductility of GFRP shear connectors decreased as the number of transverse rebars increased. Empirical equations were proposed to predict the shear resistance and validated with the test data.

1. Introduction

Fiber-reinforced polymer (FRP) materials have been widely used in civil infrastructure applications because of their excellent mechanical characteristics, such as high strength-to-weight ratio, free formability, corrosion re-sistance, and durability (Bank 2013; Hollaway 2010; Sumida and Mutsuyoshi 2008). An increasing number of new bridges have been constructed as all-FRP or hybrid FRP-concrete composite bridge structures (Nelson et al. 2014b; Cheng and Karbhari 2006; Jeong-Hun et al. 2006; Keller 2001). This increase is predominantly at-tributed to the advantageous properties of FRP compos-ites (Lee 2012). In particular, all-FRP bridge decks for rapid deck replacement with minimum traffic interfer-ence are promising applications, as observed in many

countries. However, the initial cost of all-FRP bridge decks is higher and their stiffness is lower than those of concrete decks. To overcome these disadvantages and to maximize the use of these materials, hybrid FRP-concrete composite bridge decks have been inves-tigated by a number of researchers (Nelson and Fam 2014a; Boles et al. 2015; Alnahhal and Aref 2008; Warn and Aref 2010; Bank et al. 2010; Alagusundaramoorthy et al. 2006; Cheng 2011; Hasselhoff et al. 2015; Keller et al. 2007; Gai et al. 2013; Cho et al. 2010). The advan-tages of hybrid FRP-concrete composite bridge deck systems include cost effectiveness and cross-section optimization based on the material characteristics of each component.

In steel and concrete composite structural systems, the transfer of shear forces between steel and concrete ma-terials should be ensured by arranging different types of shear connectors, such as stud connectors (Shim et al. 2004; Xue et al. 2008), perfobond rib connectors (Cândido-Martins et al. 2010; Zheng et al. 2016), T-perfobond rib connectors (Vianna et al. 2008), Y-type perfobond rib connectors (Kim et al. 2014a; Kim et al. 2013; Kim et al. 2014b), channel connectors (Baran and Topkaya 2012), angle connectors (Shariati et al. 2014), hat-shaped connectors (Kim et al. 2011), I-shaped con-nectors (Mazoz et al. 2013; Costa-Neves et al. 2013), J-hook connectors (Richard Liew and Sohel 2009), and puzzle-shape composite dowels (Lorenc et al. 2014a; Lorenc et al. 2014b). Stud shear connectors and per-fobond rib shear connectors are commonly used in steel and concrete composite structures. Studs can cause in-convenience in construction in sites to a great extent; when subjected to repeated loading, studs may suffer from fatigue failure (Wang et al. 2014; Hanswille et al. 2007b; Hanswille et al. 2007a). To overcome these

1Associate Professor, College of Civil Engineering &Hunan Provincial Key Laboratory of Structures for Wind Resistance and Vibration Control, Hunan University of Science and Technology, Xiangtan, China. *Corresponding author, E-mail: [email protected] graduate, College of Civil Engineering & Hunan Provincial Key Laboratory of Structures for Wind Resistance and Vibration Control, Hunan University of Science and Technology, Xiangtan, China. 3Associate Professor, College of Civil Engineering &Hunan Provincial Key Laboratory of Structures for Wind Resistance and Vibration Control, Hunan University of Science and Technology, Xiangtan, China. 4Professor, College of Civil Engineering & Hunan Provincial Key Laboratory of Structures for Wind Resistance and Vibration Control, Hunan University of Science and Technology, Xiangtan, China.

H. Huang, A. Li, L. Chen, C. Zeng and M. Zhu / Journal of Advanced Concrete Technology Vol. 16, 368-381, 2018 369

drawbacks, a German company developed a perfobond rib shear connector (PBL steel connector) in 1987. Since then, this connector has been extensively studied and applied in many engineering practices due to its excellent shear resistance and simple manufacturing processes. The structural performance of PBL steel connectors has been comprehensively evaluated by push-out tests and numerical analyses. The effects of design variables, such as concrete strength, number of holes, hole shape and diameter, thickness of steel plates, width and length, number of transverse rebars, and rib arrangement, have also been examined. Different analytical models have been presented to predict the shear resistance of PBL steel connectors. Various models have been established to evaluate the shear resistance of PBL steel connectors through push-out tests (Oguejiofor and Hosain 1994; Oguejiofor and Hosain 1997; Medberry and Shahrooz 2002; Veríssimo et al. 2006; Al-Darzi et al. 2007; Ahn et al. 2010; Da. C. Vianna et al. 2013). They found the three factors contributing to overall resistance: the bearing concrete resistance at connector faces, the transverse rebar at connector holes, and concrete dowels formed in rib holes. The shear resistance equations were proposed to predict the contribution of individual holes to PBL steel connector resistance (Hosaka et al. 2000; Zheng et al. 2016).

Although extensive studies on the shear connectors of steel-concrete composite structural systems have been carried out, few studies have described shear connectors of FRP-concrete composite structural systems. The main difference between steel-concrete and FRP-concrete composite structural systems is the material used for shear connectors and the connection method. As a result, the load-slip behavior of shear connectors for steel-concrete composite structural systems may vary from that of connectors for FRP-concrete composite structural systems.

In this study, two kinds of glass fiber-reinforced polymer (GFRP) shear connectors with a flat plate or a T-type rib were proposed, which were similar to PBL shear connectors in steel-concrete composite structure. Five groups of push-out specimens with GFRP shear connectors were prepared and tested. The failure modes of the specimens were observed and their load-slip curves were obtained. The effects of rib shapes, presence of holes, and existence of transverse rebars on failure mode, shear resistance, and ductility were investigated. Empirical equations were proposed to predict shear re-

sistance. These equations were then validated with the test data.

2. Push-out test

2.1 Test specimens Fifteen specimens were examined to determine the shear behavior of GFRP shear connectors with a flat plate or a T-type rib in accordance with Eurocode 4. The main variables for the experimental evaluation of a GFRP shear connector include the mechanical characteristics of concrete slabs, GFRP shear connectors, and rebars. In this study, the shape of ribs, the presence of holes, and the existence of transverse rebars were considered parame-ters to evaluate the shear capacity of GFRP shear con-nectors. Five different groups of GFRP shear connectors were prepared and tested. One group of these specimens comprised flat plate shear connectors (identified as FP), while the four remaining groups consisted of T-type rib shear connectors (denoted as TP). Table 1 summarizes the parameters of the tested specimens.

Figure 1 shows the dimensions of specimen FP-H8-R0. The concrete slabs of specimens TR-H0-R0, TR-H8-R0, TR-H4-R4, and TR-H8-R8 were of the same size but of different configurations of rib shapes, num-bers of holes, and numbers of transverse rebars in the holes. The configurations of the GFRP shear connectors are shown in Fig. 2 Each specimen was composed of one GFRP multi-box girder, two concrete slabs, and two GFRP shear connectors. The dimension of the concrete slab was 580 mm × 200 mm × 650 mm. The fabrication process is shown in Fig. 3. First, GFRP stay-in-place structural forms were prepared in a factory [Figs. 3(a) – 3(d)]. The GFRP shear connector was attached to the upper plate of the pultruded GFRP stay-in-place struc-tural form (PGSF), and this pultruded form with a shear connector could be continuously produced through pul-trusion. The surface of the PGSF was coated with silica sand to increase the bond strength at the interface. Sec-ond, two PGSFs were bonded with structural adhesive into one GFRP multi-box girder [Figs. 3(e) – 3(g)]. Fi-nally, the push-out test specimens were completed in a laboratory [Figs. 3(h) – 3(j)].

2.2 Material properties The mechanical properties of GFRP are listed in Table 2 tested in accordance with Fiber-reinforced plastics composites-Determination of tensile properties (GB/T

Table 1 Test specimens.

Hole Transverse rebar Specimens Rib shape

Number Diameter (mm) Number Diameter (mm) FP-H8-R0 flat plate 8 35 0 — TR-H0-R0 T-type rib 0 — 0 — TR-H8-R0 T-type rib 8 35 0 — TR-H4-R4 T-type rib 4 35 4 14 TR-H8-R8 T-type rib 8 35 8 14


Concrete slab

Shear connector GFRP multi-box girder

(a) Specimen 3D model

150

500

150

200 200200

P

D35

2 211 Shear connector

Hole

1-1650

580

P

connector

Hole

Shear

(b) Front view (c) Side view

2-2

5200

8080

200

580

105

200 100 100 200

connector

Hole

Shear

5

8010

80

5

35

535

5

80

100

55

35

HoleconnectorShear

plateUpper

plateVertical

plateBottom

(d) Top view (e) Cross-sectional dimensions of the GFRP component

Fig.1 Dimensions of specimen FP-H8-R0. (Unit: mm)


Hole

Flat plate

T-typed rib

(a) Shear connector in specimen FP-H8-R0 (b) Shear connector in specimen TR-H0-R0

T-typed rib

Hole

T-typed rib

Hole

Transverse rebar

(c) Shear connector in specimen TR-H8-R0 (d) Shear connector in specimen TR-H4-R4

T-typed rib

Hole

Transverse rebar

40

580

1080

5

35

535

5

10

80

100

55

Web

Flange connectorShear

plateUpper

plateVertical

plateBottom

(e) Shear connector in specimen TR-H8-R8 (f) Cross-sectional dimensions of the GFRP component

Fig.2 Configuration of GFRP shear connector. (Unit: mm)


(a) Pultrusion process (b) Pultruded GFRP stay-in-place structural form

(c) Cutting (d) Drilled holes on the web of shear connector

(e) Brushing structural adhesive on the bottom plate (f) Assembling

(g) Assembled GFRP multi-box girder (h) Arrangement of reinforcing bars and formwork

(i) Pouring concrete (j) Specimens completed

Fig.3 Fabrication processes of specimens.


1447-2005) and Fiber-reinforced plastics compos-ites-Determination of compressive properties (GB/T 1448-2005). The uniaxial compressive strength of con-crete was examined by using 150 mm cubic specimens after a 28-day curing period, determined in accordance with Standard for test method of mechanical properties on ordinary concrete (GB/T 50081-2002). Table 3 pre-sents the concrete compressive strength, tensile strength, and Young’s modulus of elasticity. Table 4 shows the material properties of transverse rebars determined in accordance with Metallic materials-Tensile testing at ambient temperature (GB/T 228-2002). 2.3 Test setup, loading procedure, and instru-mentation In this study, the push-out test proposed in Eurocode 4 was performed to evaluate the shear behavior of GFRP shear connectors. In Fig. 4, the concentric compressive load was applied monotonically by using a hydraulic jack with a capacity of 1000 kN. On the top of the assembled GFRP multi-box girder, concrete fill with a length of 75 mm was adopted to prevent local failure. A loading plate was installed on the top of the GFRP multi-box girder to achieve a uniformly distributed load between two con-nectors. In order to avoid eccentric compression of specimen, a layer of fine sand is arranged at the bottom of the specimen to ensure the uniform force of the left and right concrete slabs.

The force control was used in the initial loading stage. After a concrete crack appeared, the displacement was subsequently controlled until the test specimen com-pletely failed.

The instruments for each specimen included one load sensor and four displacement transducers. The load sensor was installed between a hydraulic jack and a ver-tical reaction frame. Four displacement transducers were symmetrically mounted at the front and back sides of the specimen and fixed at the same height. The slip between the assembled GFRP multi-box girder and the concrete slabs was measured with four displacement transducers.

3. Push-out test results

3.1 Failure modes For the non-transverse-rebar-containing specimens, such as FP-H8-R0, TR-H0-R0, and TR-H8-R0, no evident crack could be found on the GFRP connectors [Figs. 5(a) – 5(c)]. In the ultimate state, the concrete slabs involved the failure of splitting cracks. In Fig. 5(a), the failure modes of FP-H8-R0 involved the failure of concrete slabs. The initial splitting cracks appeared on the top surface of the concrete slabs. As the load increased, splitting cracks were formed below the GFRP connectors and these cracks gradually spread out toward the bottom of the concrete slabs. During the test, vertical splitting cracks could be observed on the side surface of the con-crete slabs, mostly around the GFRP connector. By con-trast, no evident deformation and crack could be detected in the GFRP shear connectors, which were intact after the test. In Figs. 5(b) and 5(c), TR-H0-R0 and TR-H8-R0 showed the same failure modes caused by the splitting crack of the concrete slabs. The initial splitting cracks appeared below the GFRP connectors and these cracks gradually spread out toward the bottom of the concrete slabs. During the test, the crack on the upper surface of the concrete slab appeared in the edge of the flange. As the load increased, several vertical cracks on the side surface could be observed. After the ultimate load was achieved, the concrete surface below the GFRP connec-tor fell off, and this observation indicated the effect of the concentrated shear forces acting at the bottom of the concrete slabs. On the contrary, no evident deformation and crack could be detected in the GFRP shear connec-tors, which were intact after the test.

For the transverse-rebar-containing specimens, such as TR-H4-R4 and TR-H8-R8, no evident crack could be found on the surface of the concrete slabs except for splitting cracks below the connectors, which was fun-damentally different from that of FP-H8-R0, TR-H0-R0, and TR-H8-R0 [Figs. 5(d) and 5(e)]. The strain meas-urements revealed no evident deformation on the trans-

Table 4 Material properties of transverse rebar. Yield strength

fy (MPa) Ultimate strength

fu (MPa) Elongation

(%) Young's modulus

Ec (GPa) 410 620 16 200

Table 2 Material properties of GFRP. Shear connector GFRP component

Longitudinal parameters flange Web Bottom plate Vertical plate Upper plate

Tensile strength (σt, MPa) 465 444 473 452 468 Tensile elastic modulus (Egt, GPa) 26.2 23.8 27.6 25.1 27.2 Compressive strength (σc, MPa) 255 235 265 245 261

Compressive elastic modulus (Egc, GPa) 11.8 9.2 12.2 10.5 11.2

Table 3 Material properties of concrete. Cube strength fc,cube (MPa)

Prism strength fc,prism (MPa)

Tensile strength ft (MPa)

Young's modulus Ec (GPa)

31.2 14.9 1.5 30.4


verse rebar. In the ultimate state, the GFRP shear con-nector was shorn off in the root with a large plastic de-formation, and the shear connector under the loading plate was crushed because of local plastic deformation. During the test, an evident sound of the tearing of a GFRP connector could be heard because of shear failure. Figures 5(d) and 5(e) describe the shearing failure modes of GFRP connectors. On the contrary, the concrete slabs were intact after the test.

3.2 Load-slip relationships In Fig. 6(a), FP-H8-R0 exhibited almost no slip in early stages because of the contribution of bonding. After the debonding between the GFRP girder and the concrete slabs occurred, FP-H8-R0 behaved nonlinearly up to the failure load. During the test, no decline phase in the load-slip curve was observed. The slip increased rapidly after the initial crack appeared on the upper surface of the concrete slabs, and splitting failure occurred suddenly.

After the test, the GFRP connectors in FP-H8-R0 were intact. The failure mode of FP-H8-R0 was the splitting crack of the concrete slabs [Fig. 5(a)].

The load-slip curves of TR-H0-R0 and TR-H8-R0 are presented in Figs. 6(b) and 6(c). The load-slip curve was divided into two phases: ascent stage and decline stage. In the ascent stage, these two specimens showed nearly no slip in the early stage before debonding occurred and then exhibited a nonlinear slip up to the ultimate load. The main difference between TR-H0-R0 and TR-H8-R0 was the variation of cracking load. In TR-H0-R0, the initial concrete crack appeared nearly at the ultimate load. In TR-H8-R0, the initial concrete cracking load was 84.9% of the ultimate load. The results indicated that the ductility of TR-H8-R0 was greater than that of TR-H0-R0. In contrast to FP-H8-R0, the two specimens with T-typed rib shear connectors manifested different behaviors in the ascent stage but showed the same failure mechanism, that is, the splitting cracking of concrete

Displacement

transducer

Load sensor

Push-out specimen

Vertical reaction frame

Hydraulic Jack

A layer of fine sand

(a) Overall view of test setup

Loadsensor

Hydraulicjack

Verticalreaction frame

Loadingplate

Supportdeck

Concreteslab

Concretefilled

Displacementtransducer

Shearconnector

(b) Loading pattern

Fig.4 Test setup and instrumentation.


slabs [Figs. 5(b) and 5(c)]. These results implied that load transfer from a GFRP girder to concrete slabs was likely ineffective if transverse rebars were absent.

Figures 6(d) and 6(e) illustrate the load-slip curves of

TR-H4-R4 and TR-H8-R8. The load-slip curve was divided into two phases: ascent stage and decline stage. In the ascent stage, these two specimens formed nearly no slip in early stages prior to debonding and then dis-

Crack on the top surface

Crack on the side surface

Crack below the position of GFRP

connector

(a) FP-H8-R0

crack on the side surface


Concrete surface fell off

(b) TR-H0-R0


crack on the side surface

Concrete surface fell off

(c) TR-H8-R0

No crack on the top surface

connector local crushed

GFRP connectortearing

GFRP connector tearing in the

root Splitting cracks below the position of connector

(d) TR-H4-R4

No crack

GFRP connector tearing

GFRP connector tearing in the

root

Splitting cracks below the position of connector

(e) TR-H8-R8

Fig.5 Failure mode of specimens.


played a nonlinear behavior up to the ultimate load. The main difference between TR-H4-R4 and TR-H8-R8 was the variation of cracking load. For TR-H4-R4, the initial concrete cracking load was 70.6% of the ultimate load. For TR-H8-R8, the initial concrete crack appeared al-most at the ultimate load. The results suggested that these specimens showed an enhanced crack resistance as the number of transverse rebar increased. The load applied to the GFRP girder was transferred to the concrete slabs

mostly by the adhesive prior to debonding. After debonding occurred, the load was transferred to the concrete slabs mainly by the concrete and the transverse rebar in the dowel holes and concrete slabs below the GFRP connectors. The behavior after debonding was nonlinear up to the ultimate load and finally failed in the decline stage because of the tearing of the GFRP con-nectors in the root. These results indicated that the ca-pacity of the GFRP connectors were insufficient to pro-

(a) Specimen FP-H8-R0 (b) Specimen TR-H0-R0

(c) Specimen TR-H8-R0 (d) Specimen TR-H4-R4

(e) Specimen TR-H8-R8

Fig.6 Load-slip curves


vide shear resistance after debonding. Thus, these specimens failed immediately because of the shear fracture of the GFRP connectors [Figs. 5(d) and 5(e)].

3.3 Effects of rib shapes In Figs. 5(a) and 5(c), FP-H8-R0 and TR-H8-R0 with different rib shapes exhibited a similar failure mode, that is, the splitting cracking of the concrete slabs. The main differences between these specimens were their crack width. The ultimate load and debonding load for FP-H8-R0 with flat plate shear connectors were 389 and 80 kN, respectively. The slip between GFRP girder and concrete slabs at the maximum load (Su) was 2.019 mm. The ultimate load and debonding load for TR-H8-R0 with T-type rib shear connectors were 465 and 99 kN, respectively. Su was 2.226 mm. The ultimate load and debonding load respectively increased by 19.5% and 23.8% when the rib shape is T-type, and Su increased by 10.3%. Figure 7 reveals the comparison results of Su, debonding load, and ultimate load between FP-H8-R0 and TR-H8-R0. In particular, the shear resistance and ductility of T-type rib shear connectors were greater than those of the flat plate shear connectors because the con-tact area between GFRP and concrete, which was formed by the concrete casting around the ribs, increased as rib shape changed. 3.4 Effects of holes on the rib TR-H0-R0 and TR-H8-R0 were characterized by the same failure mode because of the splitting crack of con-crete slabs [Figs. 5(b) and 5(c)]. The ultimate load and debonding load for TR-H0-R0 without holes were 386 and 61 kN, respectively. Su was 1.340 mm. The ultimate load and debonding load for TR-H8-R0 with eight holes were 465 and 99 kN. Su was 2.226 mm. With the holes in the shear connectors, the ultimate load and debonding load respectively increased by 20.5% and 62.3%, and Su increased by 66.1%. Figure 8 shows the comparison results of Su, debonding load, and ultimate load between TR-H0-R0 and TR-H8-R0. For the same rib shape, shear resistance and ductility were significantly improved by the holes (Fig. 8) whose presence induced an increase in

the concrete dowel area created by the concrete filling in these holes. 3.5 Effects of transverse rebars In Figs. 5(c) and 5(e), TR-H8-R8 with eight transverse rebars failed at the shear connectors rather than at the concrete slabs, and this observation was significantly different from that of TR-H8-R0 without transverse rebar. The concrete slabs were intact after loading because their shear resistance, which was caused by the concrete dowel and transverse rebar in the holes and concrete slabs be-low the position of connectors, increased due to the ex-istence of transverse rebars. In Fig. 9, the ultimate load and debonding load for TR-H8-R8 were 590 and 120 kN, respectively. Su was 1.584 mm. The ultimate load and debonding load for TR-H8-R0 were 465 and 99 kN. Su was 2.226 mm. With the transverse rebar in the shear connectors, the ultimate load and debonding load re-spectively increased by 26.9% and 21.2%, but Su de-creased by 28.8%. This finding indicated that transverse rebars significantly affect the shear resistance behavior.

Figures 5(d) and 5(e) and Table 5 provide the test results of T-type specimens with different numbers of transverse rebars. TR-H4-R4 and TR-H8-R8 showed a similar failure mode caused by the shearing failure of the connectors in the root when transverse rebars were ap-plied. In Fig. 10, the ultimate load and debonding load for TR-H8-R8 were 590 and 120 kN, respectively. Su was 1.584 mm. The ultimate load and debonding load for TR-H4-R4 were 666 and 118 kN, respectively. Su was 1.800 mm. As the number of transverse rebars increased, the ultimate load decreased by 11.4%, the debonding load remained unchanged, and Su decreased by 12.0%. Thus, the shear resistance and ductility of the GFRP shear connectors decreased as the number of transverse rebars increased because the shear area of GFRP shear connectors decreased as the number of holes increased.

4. Equations to predict the ultimate resistance of shear connectors and comparisons

Shear resistance equations for perfobond rib shear con-nectors have been examined. Nevertheless, very few

Fig. 7 Comparison of Su, debonding load and ultimate load between specimen FP-H8-R0 and TR-H8-R0.

Fig. 8 Comparison of Su, debonding load and ultimate load between specimen TR-H0-R0 and TR-H8-R0.


studies have focused on the shear connectors of FRP-concrete composite structural systems, and existing equations are unsuitable for the estimation of the shear resistance of the proposed GFRP shear connectors. Based on the failure mechanisms of push-out tests, empirical equations are presented to predict ultimate resistance and validated with the test data. 4.1 Proposed shear resistance equation when specimens failed at the concrete slabs The proposed shear resistance equation is described in Eq. (1) when the specimens without a transverse rebar failed at the concrete slabs (FP-H8-R0, TR-H0-R0, and TR-H8-R0), and this parameter depends on the concrete dowel resistance in holes and the concrete resistance by slabs.

1 2u c cV V V= + (1)

where the Vu (N) represents the shear resistance of the specimens without a transverse rebar; Vc1 (N) refers to the concrete dowel resistance by holes, as described in Eq. (2); and Vc2 (N) denotes the concrete resistance by slabs, as described in Eq. (3).

1 1 2c c tV n n A f= (2)

where n1 is the number of ribs, n2 is the number of dowel holes per rib, Ac (mm2) is the dowel areas formed in the holes, ft (N/mm2) is the tensile strength of concrete.

2 3c b h tV n f bhα α= (3)

where n3 is the number of slabs, b (mm) is the concrete slab thickness, h (mm) is the concrete slab height, αb and αh are values obtained by the regression analysis on the

experimental results related to concrete slab thickness and concrete slab height, respectively, and these values range from 0.8 to 1.0.

As mention above, Eq. (1) only can apply to the specimen without transverse rebar. However, adequate transverse rebar should be configured to avoid concrete splitting or excessive longitudinal cracks.

4.2 Proposed shear resistance equation when the specimens failed at the GFRP shear con-nectors The presented shear resistance equation is described in Eq. (4) when the specimens with a transverse rebar failed at the GFRP shear connectors in the root (TR-H4-R4, TR-H8-R8), and this parameter depends on the shear areas of the GFRP connectors in the root. Actually, the holes are generally set near the roots which significantly reduce the net web area, rather the root shear area. Moreover, test results show that specimens with a transverse rebar fail at the GFRP shear connectors in the root, rather in the holes in the web. It seems reasonable to use the full cross-section area of web to calculate the shear resistance of GFRP shear connectors. However, it is relatively safer to use net section area in engineering design.

1 2( )u sV n t l n Dτ= − (4)

where l (mm) is the length of rib; D (mm) is the diameter of holes; t (mm) is the thickness of the rib when flat plate connectors are applied, and t (mm) is the thickness of the web when T-typed rib connectors are applied, respec-tively; τs (N/mm2) is the shear strength of GFRP.

Table 6 summarizes the comparison of the tested shear resistance and the predicted shear resistance to verify the

Table 5 Summary of test results. Test specimens Pu (kN) Su (mm) Failure mode

FP-H8-R0 389 2.019 TR-H0-R0 386 1.340 TR-H8-R0 465 2.226

Concrete cracking

TR-H4-R4 666 1.800 TR-H8-R8 590 1.584 GFRP connectors tearing

Note:Pu represents the maximum load during the tests; Su represents the slip between GFRP girder and concrete slabs at the maxi-mum load.

Fig.9 Comparison of Su, debonding load and ultimate load between specimen TR-H8-R0 and TR-H8-R8.

Fig.10 Comparison of Su, debonding load and ultimate load between specimen TR-H4-R4 and TR-H8-R8.


proposed shear resistance equation [Eqs. (1) – (4)]. The predicted shear resistances were consistent with the tested shear resistances. Therefore, the proposed equa-tion not only estimates the shear resistance of flat plate connectors and T-typed rib connectors accurately within the margin of error but also enables the safe design of shear connectors. However, accounting for all of the characteristics of the presented GFRP shear connectors with only three variables provides some limitations. As such, further studies should be performed to consider the characteristics of the presented GFRP shear connectors in diverse perspectives and to develop accurate shear resistance equations.

5. Conclusions

In this study, two kinds of GFRP shear connectors with a flat plate or a T-type rib were proposed. Five groups of push-out specimens with GFRP shear connectors were prepared and tested. The effects of rib shapes, presence of holes, and existence of transverse rebars were inves-tigated. Shear resistance and ductile behavior character-istics caused by changes in variables were evaluated on the basis of experimental results. Empirical shear resis-tance equations for GFRP shear connectors were pre-sented. The following conclusions were obtained: (1) Push-out tests demonstrated that the shear strength

and ductility of T-type rib shear connectors were greater than those of flat plate shear connectors. In T-type rib shear connectors, the ultimate load and debonding load increased by 19.5% and 23.8%, respectively. Su also increased by 10.3%. Therefore, the behavior of T-type rib shear connectors was more optimum than that of flat plate shear con-nectors.

(2) The behavioral changes of the T-typed rib shear connectors were evaluated. Our results suggested that holes significantly improved shear resistance and ductility. With holes in shear connectors, the ultimate load and debonding load increased by 20.5% and 62.3%, respectively. Su also increased by 66.1%.

(3) The failure mode of the tested specimens may be determined by the existence of a transverse rebar. Without the transverse rebar, the specimens failed at the concrete slab rather than at the GFRP shear connector. With the transverse rebar, the GFRP shear connector formed and fractured in the root.

Compared with the reference specimen without transverse rebars, the ultimate load and debonding load of the specimen with eight transverse rebars increased by 26.9% and 21.2%, respectively. Con-versely, Su decreased by 28.8%. This finding indi-cated that transverse rebars significantly affected shear resistance behavior.

(4) All of the specimens exhibited a similar failure mode caused by the shearing failure of the con-nectors in the root when a transverse rebar was ap-plied. As the number of transverse rebars increased, the ultimate load decreased by 11.4%, the debond-ing load remained unchanged, and Su decreased by 12.0%. Thus, the shear resistance and ductility of GFRP shear connector decreased as the number of transverse rebars increased. Therefore, the adapt-able transverse rebar ratio was necessary to im-prove the shear resistance and ductility of GFRP shear connectors. This requirement should be sat-isfied because the shear area of GFRP shear con-nectors decreased as the number of holes increased.

(5) Empirical equations were proposed on the basis of the failure mechanisms observed through the push-out tests to predict shear resistance. These equations were then validated with the test data. The specimens without transverse rebars failed at the concrete slabs. The proposed shear resistance equations were dependent on concrete dowel re-sistance in holes and concrete resistance in slabs. The specimens with transverse rebars failed at the GFRP shear connectors in the root. The proposed shear resistance equations were associated with the shear areas of GFRP connectors in the root. The validity of the proposed equations was verified by using previously described equations and by ex-amining shear resistance.

Acknowledgments This investigation was performed in the Engineering Structure Laboratory at Hunan University of Science and Technology. This study was financially supported by the Chinese National Natural Science Foundation (Grant Nos. 51308207 and 51578236) and the Natural Science Foundation of Hunan Province (Grant No. 2018JJ3161). References Ahn, J-H., Lee, C-G., Won, J-H. and Kim, S-H., (2010).

“Shear resistance of the perfobond-rib shear connector

Table 6 Comparisons between tested shear resistance and predicted shear resistance of the proposed GFRP shear connector.

Specimens Equations determined Tested shear resistance (A) /kN

Predicted shear resistance (B) /kN Ratio (B/A)

FP-H8-R0 Vu=n1n2Acft+n3αbαhftbh wheren1=2, n2=4, n3=2, αb =0.95, αh=0.95 389 358 0.92

TR-H0-R0 Vu= n3αbαhftbh wheren3=2, αb =0.98, αh=0.98 386 366 0.95 TR-H8-R0 Vu=n1n2Acft+n3αbαhftbh wheren1=2, n2=4, n3=2, αb=1.0, αh=1.0 465 426 0.92 TR-H4-R4 Vu=n1τst(l-n2D) wheren1=2, n2=2, τs=52 666 608 0.91 TR-H8-R8 Vu=n1τst(l-n2D) wheren1=2, n2=4, τs=52 590 538 0.91


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