Experimental investigation of steel beam to RC column connection via a through-plate

Journal of Constructional Steel Research 133 (2017) 125–140

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Journal of Constructional Steel Research

Experimental investigation of steel beam to RC column connection via a through-plate Nasrin Bakhshayesh Eghbali ⁎, Seyed Rasoul Mirghaderi School of Civil Engineering, University of Tehran, P.O. Box 11365-4563, Tehran, Iran

a r t i c l e

i n f o

Article history: Received 4 August 2016 Received in revised form 13 January 2017 Accepted 10 February 2017 Keywords: Through plate Rigid shear connector RCS connection Load transfer mechanism Experimental test

a b s t r a c t Composite-framed structures provide desirable mechanical and economical features. However, methods to provide a beam to column connection with enhanced performance are currently under investigation. In the current paper, two interior connections at 3/4 scale were evaluated experimentally under cyclic lateral loading and a constant axial load on the column. In the specimens, the beams, one of I-section and the other of channel section, were connected to a vertical plate passing through the concrete column (Through Plate). To limit sliding between the steel and concrete, rigid shear connectors were employed. Steel cover plates remove any separation potential of the rigid shear connectors from the concrete while increasing the concrete strength. The load transfer occurs through three mechanisms, which include the in-plane mechanisms of the TP and side plates and the mechanism of the concrete strut. The portions of mentioned mechanisms of the entire connection moment were 65%, 20%, and 15%, respectively. By proportioning the connection components based on the presented design procedure, plastic hinges were created in the beams, and the connection components remained undamaged. The through plate involved with concrete provided a strong panel zone with elastic behavior, and the suggested connection is categorized as a fully restrained connection. The tested specimens provided permanent hysteretic diagrams without any pinching. © 2017 Elsevier Ltd. All rights reserved.

1. Introduction The method of combining steel and concrete components in a structural frame takes advantage of the appropriate features of each of these two materials for achieving better mechanical performance while decreasing the costs. A main issue in this regard is to adapt steel and concrete materials at their interface, despite the significant difference of their stiffness and strength. In the construction of building frames, concrete and steel materials have been combined in different manners, such as the steel reinforced concrete (SRC) column system, the reinforced concrete column and steel beam (RCS) frame system, and the concrete-filled tubular (CFT) column system, each having its own advantages and characteristics. On one hand, the RCS system has steel beams that provides suitable energy dissipation capacity and reduces the structure weight particularly in buildings with long spans in comparison with conventional concrete frames; on the other hand the RCS system contains reinforced concrete columns that provide the necessary lateral stiffness and reduces the costs in comparison with conventional steel frames. There are widespread research studies on the behavior of RCS frames and RCS moment connections. Sheikh et al. [1] investigated the behavior of RCS connections through 15 interior joint specimens with details ⁎ Corresponding author. E-mail address: [email protected] (N. Bakhshayesh Eghbali).

http://dx.doi.org/10.1016/j.jcsr.2017.02.007 0143-974X/© 2017 Elsevier Ltd. All rights reserved.

(including the vertical stiffener plates, steel columns, welded shear studs, and reinforcing bars) that led to significant strength increase of the connection. Cordova [2] conducted a full scale test of a 3-story, 3bay RCS frame to evaluate the design provisions and to provide information to validate models considering the strong-column weak-beam criterion, composite action of floor slab and steel beams, pre-cast column and beam-column connections. The results showed that the frame performed well in a ductile manner and did not experience any unexpected failure. Chou and Chen [3] tested a series of full-scale one-story two-bay post tensioned RCS frames to examine the performance of the frames and showed their self-centering response. Research conducted by Parra-Montesinos and Wight [4] on nine exterior connections showed that joint stirrups and steel band plates can be replaced by engineered cementitious composite material or fiber reinforced concrete. The results indicated that the RCS frame was suitable for use in a high seismic region. Liang [5] investigated experimentally two interior and two exterior space RCS connection that were designed according to a deformation-based joint design model. The specimens showed excellent seismic behavior, and the design model was effective in controlling joint deformations. The seismic response of roof level RCS connections was examined experimentally by Fargier-Gabaldón [6], and stable behavior was achieved by using the employed detail. Li et al. [7] reviewed a set of experimental and finite element research studies conducted on RCS moment frames and connections along with the development of the relevant guidelines and recommendations.

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Based on previous research studies, in connections where the beam passes through the column extendedly, there are two primary modes of failure: shear failure of the panel zone and bearing failure of concrete under high compressive stresses of the beam compression flange. To strengthen the connection against inappropriate modes of failure, different details were added to the connection and evaluated, which led to improved performance of the connection. Cheng and Chen [8] used steel band plates embedded around the column and face-bearing plates to prevent bearing failure and panel shear yielding, respectively. Alizadeh et al. [9] applied face bearing plates wider than the steel beam flanges and additional bearing plates, which were welded to the steel beam flanges on each side of the joint to increase joint bearing and shear strength. Furthermore, Alizadeh et al. [10] tested specimens with self-consolidating concrete to improve the constructability of the connections. One of the important issues concerning connections wherein the beam is cut at the column face and the column continues extendedly is to effectively mobilize the concrete at the joint area and prevent any sliding on the concrete and steel interface. Research studies, conducted on such connections, mainly carried out by the Japanese, use some details such as cover plates and horizontal stiffeners or internal diaphragms as well as extended face bearing plate with perpendicular stiffeners. The design equations for the ultimate shear strength of joints with these details along with some through-beam-type connections were offered by Nishiyama et al. [11]. Due to the complexity of composite connections and impossibility of studying the internal areas of the connection during experimental tests, there have been many finite element studies conducted. Farahmand Azar [12] studied the behavior of RCS frames based on nonlinear static analysis considering joint modeling and compared the behavior with reinforced concrete frames and showed that RCS frames had more final capacity than reinforced concrete frames. The behavior of frames with high-strength concrete columns confined by continuous compound spiral ties and steel beams was investigated by Li et al. [13,14]. Extensive parametric studies were conducted, and simplified models were proposed for hysteretic lateral load versus lateral displacement. Guo et al. [15] and Shen and Gu [16] modeled and analyzed RCS joints and studied the joint mechanisms through verified models. Habashizadeh et al. [17] developed finite element models to investigate the hysteretic behavior of RCS connections and the reliability of the finite element method. Mirghaderi and Bakhshayesh [18] proposed a new RCS connection and investigated the structural performance and stress transfer mechanisms through finite element analysis. An important issue in suggesting a connection for an RCS system is to offer a connection whose load transfer mechanism could be clearly

explained. Therefore, by quantifying the portion of load transfer paths, the design relations can be offered, as was proposed in the research studies of Deierlein et al. [19], Kuramoto and Nishiyama [20], and Mirghaderi et al. [21]. In this research, an RCS connection is proposed to improve the features of through-column type joints. To evaluate the cyclic behavior of the proposed connection, two experimental specimens were constructed at 3/4 scale and then tested. The connection load transfer mechanisms and the portion of each load transfer paths were investigated to provide adequate information for proposing the design procedure. The suggested connection uses a vertical plate (named the through plate (TP)) that passes through the column at the joint area in both directions. The TP connect steel beam to the reinforced concrete column and is connected to the cover plates, surrounding the column at the joint area. Furthermore, to provide sufficient connection between the cover plates and TP with the concrete column, rigid shear connectors were used. While rigid shear connectors provide sufficient strength, they are able to limit sliding between steel plates and concrete more effectively than flexible shear connectors, which require primary relative displacement for being mobilized [22]. 2. Research importance Some of concerns related to RCS connections are the pinching of hysteretic responses and the complexity of the joint area. The pinching of the hysteretic responses generally occurs via local damages and concrete crushing. The complexity of the joint area and associated problems, such as reinforcement arrangement and concrete placement, typically lead to the use of special types of concrete. In the proposed connection, on one hand, the beam demands are transferred to the column by in-plane action of the TP, and local damages caused by out-of-plane forces are prevented. On the other hand, by cutting the beams at the column face and merely passing a plate through the joint area, the reinforced concrete column continues extendedly in this area, and by providing a continuous access inside the column, an appropriate arrangement of the column reinforcement is provided, while simplifying the concrete placement. Moreover, the majority of the developed forces in the plates of the joint area are transferred to the concrete via shear connectors; thus, local damages due to reliance of the TP edges to the concrete and bearing failure are prevented. As was expected, the results showed an excellent cyclic behavior without any pinching and strength and stiffness degradation. Note that the appropriate behavior of the connection is achieved by using regular concrete with 40 MPa of compressive strength.

Fig. 1. Connection configuration: (a) The first specimen (TPI), (b) The second specimen (TPU).

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Fig. 2. Connection assembly: (a) The first specimen (TPI), (b) The second specimen (TPU).

3. Characteristics of the specimens The conducted study involves experimental evaluation of two interior connections of a moment RCS frame. In the first specimen (TPI), the beam includes an I-section. To show the capability of using the beam with different sections, the I-section beam was replaced with two channel section ones in the second specimen (TPU). In both specimens, the TP at the joint area was made cruciform to provide the possibility of assembling the beams in two perpendicular directions. Passing the TP in the perpendicular direction causes separation in its surrounding concrete in the panel zone. Therefore, its effect on the composed concrete diagonal strut will be considered. Cover plates surround the column at the joint area and their edges are not involved with the concrete because of the low strength of concrete cover to tolerate compression forces from the edges of the cover plates. Rigid shear connectors were used to transfer shear forces and to prevent sliding of the steel plates relative to the concrete. Rigid shear connectors in the concrete confined by cover plates can transfer high shear forces to the concrete [21]. Steel cover plates removes any separation potential of rigid shear connectors from the concrete, while increasing the concrete strength. Research studies show that the strength of confined concrete with steel cover plates can considerably increase [23,24]. Although the upper and lower edges of the TP in contact with concrete resists sliding, in the case of not using shear connectors, the high level of compression stresses on the TP edges lead to local damage and concrete crushing in these areas. In the tested specimens, although beams are connected to the column in one direction, the required shear connectors for an orthogonal connection are attached. In fact, half of the employed shear connectors transfer 80% of total shear forces on the steel concrete interface. The purpose of using shear connectors in the perpendicular direction was to consider their reducing effect on the concrete core of the joint area and considering their contribution in the load transfer.

throughout the required length for groove welds based on Fig. 3 (b). The connection of the beam web to the TP is provided by fillet welds. It is notable that the beam end is located 20 mm from the column face and no force is directly transferred from the beam to the column. The connection of the rigid shear connectors to the TPs and the cover plates are provided by fillet welds all around. Afterwards, the TPs are connected in cruciform. The cover plates are then connected to the TPs via groove welds. The cover plates are also welded to one another. The constructed steel box is a part of the formwork of the column in the joint area. Because the joint area is surrounded by the cover plates, there is no necessity for stirrups inside the cover plates.

4. Construction process The overall shape of the specimens is depicted in Fig. 1(a) and (b). Furthermore, the connections components in details are proposed in Fig. 2(a) and (b). In specimen TPI, the beam web is omitted throughout the required length for the connection and the TP is placed in the aligned gaps of the flanges. The connection of the beam flanges to the TP is provided by groove welds after beveling the beam flanges in this area according to Fig. 3 (a). Fig. 2 (b) illustrates only those components of specimen TPU which are different from specimen TPI. In specimen TPU, the beams include two channel sections, in which the corners of the beam section on the intersection of the web and flanges are beveled

Fig. 3. Beveling of the beam flanges for groove welds to the TP: (a) The first specimen, (b) The second specimen.

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Fig. 4. (a) The forces acted on the joint area, (b) Weld reaction forces in the first specimen, (c) Weld reaction forces in the second specimen.

5. Philosophy of using the proposed details

6. Joint resisting mechanisms

The panel zone of a beam to column connection should provide a proper base for changing the direction of beam forces. In RCS connections, in addition to providing sufficient strength and stiffness against induced shear forces, an adequate connection between two types of materials should exist. In the proposed details, the steel plates of the joint area including the TP and side plates increase the shear strength of the panel zone and are engaged with concrete via shear connectors. Furthermore, cover plates mobilize the concrete core for carrying a part of shear stresses of the panel zone.

Fig. 4 (a) shows the applied forces on the joint area. The majority of the moment resulting from the beam plastic hinges turns into tension and compression forces in the beam flanges. Afterwards, these forces are transferred to the TP through the longitudinal welding lines of the flanges. Fig. 4 (b) demonstrates the welding lines of the flanges and web connection plate in specimen TPI. In specimen TPU, the beam shear forces and the portion of the beam web moment are transferred directly via the welding line of the web (Fig. 4 (c)). The applied forces in the TP plane want to rotate this plate relative to the column. The

Fig. 5. External forces and their reactions on the: (a) TP, (b) Face plate, (c) Side plate, (d) concrete panel.

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Fig. 6. Resisting mechanisms of the proposed connection: (a) The column section, (b) Spring model.

tensile stresses in the TP along the beam tensile flange decrease while moving into the column. The reason behind the sudden drop of the tension stresses in the TP is that a portion of the tensile forces are transferred to the cover plates on the front and back sides of the column (Face plates). In fact, according to the stiffness of the connection components, the total demands are distributed between the TP and face plate. The induced moment in the TP plane, inside the column, causes shear stresses that are equal to a couple of shear forces on two opposite sides of the plate (Fig. 5 (a)). Some portion of these shear forces are transferred via the shear connectors installed on the TP as well as the upper and lower edges of the TP in contact with concrete. The remaining forces are transferred to the face plate via weld connection, finally being transferred to the column through the shear connectors of the face plate. The induced moment in the face plate creates out-of-plane tensile and compression forces on its surface (Fig. 5(b)). Tensile forces mobilize the cover plates parallel to the beams (Side plates). The mobilized moment in the side plates can be replaced by a couple of shear forces (Fig. 5(c)). The shear connectors on the side plates transfer a portion of these shear forces to the column, while the remaining forces are transferred to the column through the shear connectors of the face plate. The compression forces are applied on the concrete surface. These compression forces along with the compression forces from the column form a concrete diagonal strut (Fig. 5(d)). Consequently, the located forces in the TP at the column face encounter three parallel mechanisms, as shown in Fig. 6(a): the TP mechanism, the side plates mechanism, and the concrete strut mechanism.

The stiffness of each component of the load paths can be displayed as a spring, as depicted in Fig. 6(b). 7. Contribution of the resisting mechanisms in the load transfer The distribution of connection moment (Ma) occurs between the TP and the face plate, based on the stiffness of each path. The face plate portion of the moment mobilizes the side plates as well as the concrete strut mechanism. Consequently, the total moment of the connection can be computed as follows: Ma ¼ M TP þ MSP þ M DS

ð1Þ

where MTP, MSP, andMDSare the induced moment in the TP, side plates, and concrete panel, respectively. As described in Section 6, the induced moment in the TP and the side plates can be replaced by a couple of shear forces on two opposite sides of the plate, in accordance to Fig. 5 (a) and (C). The shear strength of the plates and the shear connectors involved with concrete should tolerate these shear forces, as follows, VTP ¼

2MTP 0 þ Vb ¼ V PT þ n:STP :d =dc dc

ð2Þ

where VTP is the shear force of the TP mechanism; Vb is the beam shear force; VPT is the shear forces of the TP; n is the number of the shear connectors on one half of the TP (See Fig. 7); and STP is the shear demand of

Fig. 7. Induced forces in the shear connectors.

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the moment of the TP to the side plates were obtained 3.2 and 3.43 for Specimens TPI and TPU, respectively. 7.2. Contribution of the concrete diagonal strut The portion of the concrete strut is determined by three equations including force and moment equilibrium and compatibility of displacements, with the assumption of linear strains in the face plate in accordance with Fig. 8, presented as below, 8 > > > > > <

∑F y ¼ 0 → f s ¼ f p

D:x 2t s ðB−xÞ

(6)

2 1 f t s ðB−xÞ2 þ f p Dx2 ¼ M FP ¼ M a −MTP ∑M O ¼ 0 → > 3 s 3 > > > > : Deformation Compatibility→εs ¼ B−x → f ¼ f p Es B−x x εc Ec x s   Ec 6; 8⇒ D −2 t s x2 þ 4t s B:x−2t s B2 ¼ 0 Es

(7) (8)

ð9Þ

Fig. 8. Equilibrium of the applied forces on the face plate.

each shear connector on the TP. d′is the distance between the center line of the shear connectors on the through and also the side plates (in accordance to Fig. 7) and dc is the column width. Moreover, a relation for the side plates can be derived in similar way as follows, V SP

The distance of the plate rotation center from the compression end of the plate (x) can be determined from Eq. (9). By substituting x in Eq. (6), the ratio of the moment of the concrete strut to the side plates mechanism (including the side plates and their shear connectors) can be calculated, which was 0.77 for the specimens. 7.3. Contribution of the shear connectors

2MSP 0 ¼ ¼ V PS þ m:SSP :d =dc dc

ð3Þ

where VSP is the shear force of the side plates mechanism; VPS is the shear force of the side plates; m is the number of the shear connectors on a half of the side plates (in accordance to Fig. 7), and SSP is the shear demand of each shear connector on the side plates. 7.1. Contribution of the TP and the side plates Determining the portion of the TP and the side plates of the joint demands depends on the thickness and height of these plates and the face plate, the column dimensions, the concrete strength and the shear connectors' resistance. Therefore, determining the portions for various connections requires accompanied numerical investigations. Here, however, the portion of the TP has been determined based on the ratio of the maximum recorded shear strain in the TP (γTP) to the maximum recorded shear strain in the side plates (γSP), from the experimental tests. Based on the results presented in Section 9.4, the ratio of the maximum shear strain of the TP to the side plate (α) was 3.09 and 3.11 in specimens TPI and TPU, respectively.

As aforementioned, strength of the plates mechanisms is in fact influenced by the shear strength of the plates and the strength of shear connectors. Therefore, it is required that the shear connectors portion in load transfer to the concrete should be determined at first. In fact the TP, the face plate and the side plates are medium components for load transfer to the column and each path ends to the shear connectors. The activation of the shear connectors on the TP and the side plates rely on the shear deformations of the mentioned plates. Therefore, it can be said that the transferred force to the concrete by the TP's shear connectors are α-fold the force transferred by the side plates' shear connectors (α was 3.09 and 3.11 in Specimen TPI and TPU, respectively). The moment transferred by all shear connectors along with the concrete strut is equal to the entire connection moment. Moreover, the ratio of the shear connectors' demand on the face plate to the TP could be determined based on the compatibility of displacement. As a result, the three equations can be written as below: STP ¼ α:SSP 0

γ TP ¼ α:γ SP V PT ¼ α:

ð4Þ

V UPT :V PS V UPS

ð5Þ

VUPT and VUPS are the shear capacity of the TP and the side plates, respectively. Accordingly, the ratio of the shear force and consequently,

ð10Þ 0

n:STP :d þ 2m:SSP :d þ p:SFP :dc þ M DS ¼ Ma

STP ¼

ð11Þ

0

d SFP dc

ð12Þ

where SFP is the shear demand of each shear connector and P is their number, installed on the face plate.

Table 1 The contribution of the shear connectors and concrete strut of the total moment. Shcs on the FP Specimen TPI TPU

mFP (kN·m) 198.16 206.73

Shcs on the SP Shear demand (kN) 9774.25 10,196.56

mSP (kN·m) 15.24 15.9

Shcs on the TP Shear demand (kN) 1503.73 1568.7

mTP (kN·m) 45.73 47.71

Concrete strut Shear demand (kN) 4511.19 4706.11

MDS (kN·m) 45.73 47.71

Shear demand (kN) 3759.33 4189.50

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Fig. 9. Calculation of the column to beam capacity ratio.

Based on the three above equations, the portion of each shear connector of the total moment is as below: SSp ¼ STP

mSP ¼

 ∝

0

¼ SFP :d

. ∝:dc

02

¼

1 0

0

2

n:∝:d þ 2m:d þ p:∝:dc =d

0

ðM a −M DS Þ

mTP mFP :d 1 ¼ ¼ ðMa −MDS Þ 2 2 02 ∝ n:∝ þ 2m þ p:∝:dc =d ∝:dc

ð13Þ

the cover plates. As illustrated in Fig. 9, the column moment at the intersection of the beam and column centerline was calculated based on enlarging the column moment capacity on the top and bottom of the cover plates, with considering the axial force based on ACI318 Code [25]. The moment ratios were 1.16 and 1.25 for specimens TPI and TPU, respectively, which show that the column and beam capacity has been selected closely.

ð14Þ

where mTP, mSP, and mFP are the transferred moment by each shear connector, installed on the TP, side plates, and face plate, respectively. In the test specimens, 15%, 5%, and 65% of the entire connection moment are transferred to the column by shear connectors, installed on the TP, side plates, and face plate, respectively. By means of Eqs. (2) and (8), the portion of each connection component of the joint total moment can be obtained. The portion of the TP mechanism (MTP), the side plate mechanism (MSP), and the concrete strut (MDS) of the entire connection moment is 65%, 20%, and 15%, respectively. The shear connectors located on the face plate, because of their position at farther distance in relation to the neutral axis of the column, provide more participation in transferring the shear forces to the column than the shear connectors, located on the TP and the side plates. Therefore, to simplify the connection detail, in an optimal design, a number of shear connectors on the TP can be replaced by less number of shear connectors on the face plate. The contribution of the shear connectors and concrete strut in the entire moment of the connections is presented in Table 1. 8. Connection design The basic rule for the design of the specimens was strong-column weak-beam criteria, with the aim of forming plastic hinges in the beam. The connection details are designed based on the expected forces of the plastic hinges and they are proportioned according to the contribution of resisting mechanisms, mentioned in Section 7. Therefore, it is expected that the beams contains plastic hinges and the damages in the column and connection are limited. Since the joint area is reinforced with cover plates and TP, the most likely location for damage in the column is the upper or lower area of

8.1. Design of the TP The net section of the TP at the end of the beam (Section A-A in Fig. 4(a)) should be checked against the generated demands. Therefore, the plastic moment of the TP (Mpp) is compared with the moment demand at section A-A (Ma). 2

Mpp ¼ t p hp f yp =4NM a

ð15Þ

where tp, hp and fyp are the thickness, the height at section A-A and the yield strength of the TP, respectively. Ma = Mpb + Vb . a where Mpb is the beam plastic moment, Vb is the beam shear capacity, and a is the distance between the hinge location and section A-A. The ratio of the existing moment to the moment capacity of the TP for the specimens TPI and TPU was 0.91 and 0.95, respectively, as presented in Table 2. 8.2. Design of the beam to the TP connection In specimen TPI, the tensile forces of the beam flanges (T = Mbf/db), resulting from the moment demand at Section B-B (see Fig. 4(a)) are transferred along the welding lines of the flanges to the TP, in which

Table 2 Design of the TP at the beam end. TP at the beam end Thickness Height at the Specimen Mpb (kN·m) (mm) beam end (mm) TPI 246.96 20 500 TPU 252.3 20 500

Moment capacity (kN·m) 332.50 332.50

Moment demand (kN·m) 301.46 314.49

Moment ratio 0.91 0.95

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Table 3 Design summary of the connection components. TP in the panel zone Spec. Ma (kN·m) TPI TPU

304.87 318.05

MPT Shear (kN·m) capacity (kN) 152.43 1483.47 159.02 1506.7

SPs in the panel zone Shear demand (kN) 1270.28 1445.64

Shear ratio 0.86 0.96

MPS Shear (kN·m) capacity (kN) 45.73 1711.6 47.71 1711.6

FP Shear demand (kN) 383.2 436.11

Shear ratio 0.22 0.25

MPT Shear (kN·m) capacity (kN) 152.43 2324.9 159.02 2324.9

Shear demand (kN) 1325.51 1514.48

Shear ratio 0.57 0.65

Mbf is the beam flanges portion of the moment at section B-B and db is the beam height. In specimen TPU, the web connection of the beam is located at the end of the beam length (Fig. 4(c)). Therefore, the tensile forces of the beam flanges (T) are calculated based on the moment demand at Section A-A, which must be transferred along the welding lines of the beam flanges. The required thickness of fillet welds to transfer the web demands is determined based on the beam shear and the web moment.

9. Experimental program

8.3. Design of the panel zone

The set up configuration is presented in Fig. 10. The specimens are constructed cruciform that are extracted from a moment frame with 4 m beam length and 3 m column height. Therefore, for simulating the boundary conditions in the experimental tests, a real hinge was used at the column base and roller supports were used at the beam ends. The roller support allows free movement in the direction of the lateral loading in addition to free rotation. To retain the specimens in the loading plane, the column was supported above the specimen at the height of lateral load application. In specimen TPU, two lateral supports were set 0.7 m from the beam supports. The maximum spacing of the beam lateral supports was 1.8 m, based on the AISC seismic provisions. By considering the lateral support of the column, the necessary beam lateral supports were provided. However, because there was not any lateral support at the location of the beam plastic hinges, increasing lateral-torsional buckling led to strength degradation during 7% story drift. In specimen TPI, two lateral supports were added at the location of the beam plastic hinges.

The shear strength of the panel zone (Vn) is provided by three components, including 1) the TP and its shear connectors, VTP, 2) the side plates and their shear connectors, VSP, and 3) the concrete panel, VCS. Each component should tolerate the induced shear forces separately. The strength of the panel zone is mainly provided by the TP. The shear capacity of the TP (Vu) is calculated based on Ref. [26] by considering the strength of the corner areas as below:   3:46ω2 V u ¼ 0:577 f yp dc t p 1 þ dc db

ð16Þ

where ω is the width of the corner element of the TP. The shear demand of the panel zone is as the follows, V PZ ¼

V PZ

2ðMTP −n:mTP Þ db  Vu

This research includes two cyclic tests on a new RCS connection by the TP. The compressive strengths of the concrete were 42 MPa and 40 MPa for the first and second specimens, respectively. The mechanical features of the steel materials resulting from tensile tests are presented in Table 4. 9.1. Test set up

ð17Þ

was 0.85 for both specimens that provides elastic behavior for

the panel zone of the specimens all over the test. 8.4. Design of the face plate As aforementioned, a portion of the shear forces induced in the TP, is transferred to the concrete column via its shear connectors (n .STP) and the rest of the shear forces (VPT) is transferred to the face plate via the groove welds (in accordance with Fig. 5(b)). Therefore, the section of the face plate should resist mentioned shear forces as well as the outof-plane moment of the face plate. Deign of the connection components is summarized in Table 3. 8.5. Design of the shear connectors The number of shear connectors on each plate of the joint area is determined in a way that the strength of each shear connector will be more than the existing shear demand. The strength of the shear connectors has been determined based on the tests, conducted in ref. [20] and was used in the design of both specimens. Note that, in the case of using shear connectors with higher capacity, which is provided by increasing penetrating depth of the shear connector or using concrete with higher strength in the connection area, the number of shear connectors can be decreased. Furthermore, with additional studies and conducting cyclic tests for confined rigid shear connectors, design equation for shear connector can be modified.

9.2. Test specimens The section of the reinforced concrete column was 400 mm × 400 mm with 12 Φ18 steel bars for both specimens. In the first specimen, the beam included an I-section, with flange and web dimensions of 130 mm × 15 mm and 230 mm × 10 mm, respectively. The beam was selected a double section of UNP220 for the second specimen. To restrict lateral buckling of each beam separately, two channel beams were stitched at 65 cm intervals. Figs. 11 and 12 demonstrate the details of specimens TPI and TPU, respectively. To transfer shear forces from the cover plates and TP to the concrete, rigid shear connectors with the dimension of 70 mm × 30 mm were installed on four dual rows on each surface. Table 4 Mechanical features of the steel materials. Test spec.

Member

Yield strength (MPa)

Tensile strength (MPa)

Elongation (%)

1 1 2 2 1,2

Beam flange (I) Beam web (I) Beam flange (UNP220) Beam web (UNP220) Cover plate (thickness = 12 mm) TP (thickness = 20

329.6 243.5 321 339 309.1

474.5 448 463 471 414.1

30.2 41 27.6 30.1 35.6

256.2

446.8

33.4

544.4

687.4

16

426.8

660.4

19

1,2 1,2 1,2

mm) Bending reinforcement φ18 bar Stirrupφ10 bar

N. Bakhshayesh Eghbali, S.R. Mirghaderi / Journal of Constructional Steel Research 133 (2017) 125–140

Fig. 10. Test setup configuration (dimensions are in mm).

Fig. 11. Specifications and dimensions of the first specimen (dimensions are in mm).

133

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Fig. 12. Specifications and dimensions of the second specimen (dimensions are in mm).

connection strength, stiffness and ductility. The level of axial load was controlled via a load cell, located above the column.

9.3. Loading pattern The specimens were subjected to lateral displacement via two compressive actuators on the column tip. The lateral displacement included two cycles at 0.25%, 0.375%, 0.5%, 0.75%, 1%, 1.5%, 2%, 3%, 4%, 5%, 6%, 7% and 8% total story drift angles. To record the amounts of lateral loads, a load cell was placed on both side of the column between the load actuator and the column. An axial load of 500 kN was exerted by four prestressed bars out of the column to consider the effect of axial load on the

9.4. Measurement technology The lateral displacements in the load direction were measured by two LVDTs on both sides of the column tip. LVDT Nos.1 and 2 were connected diagonally to the four corners of the TP within the panel zone area in accordance with Figs. 13 and 14 to measure the shear

Fig. 13. Instrumentation of the first specimen (dimensions are in mm).

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Fig. 14. Instrumentation of the second specimen (dimensions are in mm).

deformations. LVDT Nos.3 and 4 were installed to determine the relative rotation between beam and column. The sliding between cover plates and concrete was measured by LVDT No.5. In addition, some linear and rosette strain gauges were installed on the different locations of the specimens in accordance with Figs. 13 and 14. 10. Test results 10.1. Failure mode Fig. 15 illustrates specimen TPI after 8% story drift. Because the specimens were designed based on strong-column weak-beam criteria, the plastic hinges were formed in the beams after the TP and the joint area stayed undamaged. Fig. 16 (a) and (b) present the cyclic lateral loads versus story drifts for specimens TPI and TPU, respectively. Specimen TPI showed a stable cyclic behavior up to 8% story drift without any

Fig. 15. Failure mode of the first specimen (TPI).

Fig. 16. Cyclic lateral loads versus story drift: (a) The first specimen (TPI), (b) The second specimen (TPU).

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strength and stiffness degradation. Specimen TPU did not exhibit any strength degradation up to 6% story drift; afterwards, it experienced a 17% of strength reduction up to 8% story drift. Hence, both specimens satisfied the AISC seismic provision requirements for qualifying as a special moment connection. The strength degradation, demonstrated in the hysteresis diagram of the second specimen, occurred after lateral-torsional buckling at 6% story drift.

observed in the white wash coating of the TP after the beam in the net section of the plate. Lateral-torsional buckling occurred at 6% story drift, which increased rapidly. At 7% story drift, as local deformations increased in the beams, some damage was observed at the end of the weld connection of the beam to the TP and the test was terminated. Fig. 18 (a)–(d) show the test observations from specimen SPU at 8% drift.

10.2. Test events

Fig. 19 (a) and (b) illustrate the normalized strain in the beam flanges and web versus the story drifts for specimen TPI and TPU, respectively. The normalized strain resulted by dividing the recorded strain via the strain gauge to the steel yield strain. The yielding started in beam flanges immediately after the TP at the location of strain gauges Top2 and Bot.2 at 1.5% story drift for both specimens. The measured strains of strain gauges Top1 and Bot.1 show that plastic strains did not spread significantly in the connection zone. In the first specimen, the plastic hinge was gradually concentrated in a place after the web connection plate. As shown in Fig. 19 (b), the strains decreased after 6% story drift in the second specimen. The decrease of the strains was due to the lateral-torsional buckling of the specimen and consequent changes in the local deformations of the beam flanges. Based on the measured strains in the middle of the beam height, plastic strains expanded throughout the beam web at the location of the plastic hinge in both specimens at the 3% drift. The maximum induced moment at the plastic hinge location was 247.1 and 226 kN·m in the process of loading for specimens TPI and TPU, respectively, which was calculated based on the maximum imposed lateral load on the specimens. The ratio of the specimen strength

10.2.1. Specimen TPI Yielding initiated at 1.5% story drift in the beam flanges after the TP. Next, some slight flexural cracks appeared in the concrete column at the top and bottom of the cover plates during the first cycle of 2% story drift and expanded during 3% story drift. At 3% story drift, the local buckling of the beam flanges occurred. At 5% story drift, some cracks were observed in the white wash coating of the TP in the corner areas. At 7% story drift, lateral-torsional buckling of the left and right beams occurred. The test continued up to 8% story drift; no crack or damage occurred in the welds until the end of the test. Fig. 17 (a)–(d) illustrate the test observations related to specimen TPI at 8% story drift. 10.2.2. Specimen TPU Plastic strains started in the beam flanges immediately after the TP at 1.5% story drift. Some slight flexural cracks appeared in the concrete column on the top and bottom of the cover plates at 2% story drift, which expanded slightly until 4% story drift. At 4% story drift, the local buckling started in the beam flanges. During 5% story drift, some cracks were

10.3. Beam web and flanges

Fig. 17. Specimen TPI at 8% story drift. (a) Formation of beam plastic hinges and concrete crack propagation, (b) Plastic deformations in the TP, (c) Web and flange connection of the beam to the TP, (d) Beam lateral buckling.

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Fig. 18. Specimen TPU at 8% story drift. (a) Formation of the beam plastic hinges and concrete crack propagation, (b) Plastic deformations in the TP, (c) Weld connection of the beam flanges to the TP, (d) Beam lateral buckling.

to the nominal plastic hinge moment obtained in accordance with the yield stress of the beam steel materials was approximately 1.23 and 1.31. This increase was due to the strain hardening phenomenon in the beam steel materials, which occurred at high plastic strains. 10.4. Panel zone In the proposed connection, the panel zone includes the TP, the side plates, and the concrete of the joint area. To record the strain changes in the panel zone and determine the portion of its components, a rosette strain gauge was installed on the TP and the side plate. Fig. 20 illustrates the normalized Mises strains against story drift in the TP and the side plate. As shown in Fig. 20, the TP and the side plate of both specimens remained elastic during the loading process. Furthermore, the maximum plastic strain in the TP was approximately 3.09 and 3.11 times greater than the maximum plastic strain in the side plate for specimen TPI and TPU, respectively. These ratios indicate the different participation level of the TP and side plates in the shear stresses of the panel zone. Fig. 21 (a) and (b) show normalized strain along the height of the TP versus the story drift for specimens TPI and TPU, respectively. The pasted strain gauges on the TP in adjacent to the beam flanges showed the highest amount. The strains gradually decreased by moving towards the plate edges. In specimen TPI, two strain gauges were installed between the beam flanges on the TP. According to Fig. 21 (a), the measured strains in the location of strain gauge TP5 show that the plastic strains expanded throughout the TP depth at 4% story drift. Note that, even though the plastic strains occurred in the TP, plasticity expanded in the beam with a faster growth, leading to the formation of plastic hinges in this area.

Fig. 22 (a) and (b) demonstrate the lateral load against the story drift caused by shear deformation of the TP for specimens TPI and TPU, respectively. In both specimens, the TP behavior was relatively elastic and the maximum participation of the TP in the story drift was 0.28 and 0.22 at 8% story drift for specimens TPI and TPU, respectively.

10.5. Cover plates To study the amount of strains in the cover plates, two strain gauges were pasted on the cover plates in both specimens in accordance with Figs. 13 and 14. Based on Fig. 23 (a) and (b), the cover plates stayed elastic in both specimens. Cover plates, while providing the confinement of the concrete at the joint area, provide the necessary lateral support for the shear connectors. The location of the strain gauges on the cover plates mainly reflects the induced strains by shear connectors in the side plate and face plate.

10.6. Evaluation of sliding To measure sliding between the cover plates and the concrete, an LVDT was installed at the steel-concrete interface, and the recorded displacement versus the story drifts for both specimens is presented in Fig. 24. The amount of sliding at the maximum lateral load reaches 0.9 and 0.72 mm for specimens TPI and TPU, respectively. As observed in hysteresis diagrams, this small amount of sliding had no effect on the overall behavior of the connection.

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Fig. 21. Normalized strains in the TP height versus story drift: (a) The first specimen (TPI), (b) The second specimen (TPU) (see Figs. 13 and 14 for numbering of the strain gauges). Fig. 19. Normalized strains of the beam web and flanges versus story drift: (a) The first specimen (TPI), (b) The second specimen (TPU) (see Figs. 13 and 14 for numbering of the strain gauges).

10.7. Column reinforcements To study the reinforcement behavior, the strains in the corner and middle reinforcements were recorded at a distance of 15 cm from the last stirrups, adjacent to the joint area (10 cm inside the cover plates) in accordance with Fig. 25 (a). The normalized strain of the corner and

middle reinforcements versus the story drift is given in Fig. 25 (b). The amounts of strains reached 0.82 and 0.74 of the yield strain in the first and second specimens, respectively, indicating that the columns were optimally designed. 10.8. Connection rigidity One of the important criteria for categorizing a connection is the connection rigidity according to the AISC Specification [27]. To study the proposed connection rigidity, the moment against relative beamcolumn rotation at service loads is presented in Fig. 26 (a) and (b). The AISC Specification introduces the secant stiffness of a connection (KS) as the moment to the relative rotation ratio at service loads, which is used as a criterion to evaluate the rigidity of a connection. Secant stiffness is compared with the ratio of bending rigidity to the length of the beam. For fully restrained connections, the secant stiffness should exceed 20 times more than this ratio. As seen in Fig. 26 (a) and (b), the tested connections provided adequate stiffness at service loads and they can be considered rigid connections in accordance with the AISC Specification. 11. Conclusions

Fig. 20. Strain response in the panel-zone for both specimens.

In this paper, an RCS moment connection was introduced that uses the TP to transfer the induced demands. Next, two interior connections in 3/4 scale were evaluated experimentally under cyclic lateral loading and a constant axial load on the column. The portion of the load paths

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Fig. 23. Normalized strains in the cover plates height versus story drifts: (a) The first specimen (TPI), (b) The second specimen (TPU) (see Figs. 13 and14 for numbering of the strain gauges).

• The evaluation of the connection rigidity, based on the AISC Specification, demonstrates that the suggested connection can be categorized as a fully restrained connection. Fig. 22. Cyclic lateral loads versus story drift developed by shear deformation of the TP: (a) The first specimen (TPI), (b) The second specimen (TPU).

along with the designing procedure was presented. Based on the results, the following can be concluded: • A characteristic of the tested connection is that all demands of the beam plastic hinges turn to in-plane forces in the TP plane. The TP increased the shear capacity of the joint area to tolerate the shear demands of the panel zone. • In general, the beam demands are transferred to the concrete column through three mechanisms: the in-plane mechanisms of the TP and side plates and the mechanism of the concrete strut. • Each load transfer paths end to the shear connectors that transfer the induced forces to the column. An insignificant amount of measured sliding at the steel-concrete interface (less than 1 mm) throughout the loading process indicates that the rigid shear connectors possess adequate stiffness and strength to transfer the loads without considerable sliding. • Plastic strains started in the beams immediately after the TP at 1.5% story drift for both specimens, and the plastic hinges expanded in the beams. By proportioning the connection components based on the presented design procedure, the joint area remained undamaged throughout the test, and only slight cracks occurred in the concrete column. • Specimen TPI and TPU demonstrated a stable hysteretic behavior without pinching until 8% story drift; this result highly exceeded the AISC seismic provision requirements for special moment-resisting frames.

Acknowledgements The authors would like to appreciate Padena Hoor Company for providing materials and construction costs and the Building and Housing Research Center (BHRC) of Iran for conducting the experimental work. Furthermore, the authors wish to acknowledge Mr. M.M. Ahmadi for valuable assistance with this project.

Fig. 24. Relative displacement at the steel and concrete interface versus story drift for both specimens.

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Fig. 25. (a) Instrumentation of the corner and middle reinforcements, (b) Normalized strains of the column reinforcements versus story drift.

Fig. 26. Beam moment versus relative beam-column rotation: (a) The first specimen, (b) The second specimen.

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