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Behavior of welded CFT column to H-beam connections with external stiffeners
Engineering Structures 26 (2004) 1877–1887 www.elsevier.com/locate/engstruct
Behavior of welded CFT column to H-beam connections with external stiffeners Kyung-Jae Shin a, Young-Ju Kim b, Young-Suk Oh c, Tae-Sup Moon b, a
b
Department of Architectural Engineering, Hannam University, 133 Ojung-dong taeduk-gu, Taejon 306-791, South Korea Department of Architectural Engineering, Hanyang University, 17 Haengdang-dong, Seongdong-gu, Seoul 133-791, South Korea c Department of Architectural Engineering, Taejon University, 96-3 Yongun-dong, Tong-gu, Taejon 300-716, South Korea Received 5 September 2003; received in revised form 21 June 2004; accepted 23 June 2004
Abstract This paper focuses on the experimental and analytical behavior of concrete filled tubular (CFT) column to H-beam welded moment connections with external T-stiffeners. Six full-scale specimens were tested cyclically. All specimens were designed and manufactured using the following main test parameters: (a) the strength ratio of the horizontal stiffener to the beam flange; and (b) the strength ratio of the vertical stiffener to the beam flange. Three types of failure modes, horizontal stiffeners failure, vertical stiffener failure and beam failure, were obtained. The connections reinforced with T-stiffeners having 130% of strength to beam flanges showed stable hysteretic behavior and good ductility. The ABAQUS finite element package was used to simulate the experimental behavior. The results obtained from the finite element were evaluated by comparing the load–displacement responses and the potential of failure modes. # 2004 Elsevier Ltd. All rights reserved. Keywords: Concrete filled tube (CFT); T-stiffener; Moment-resisting frame; Plastic rotation
1. Introduction Steel–concrete composite materials have been widely used in framing systems for high-rise buildings. One particular type of composite construction is concrete filled steel tube (CFT). CFT columns have been found to have advantages especially for the column loaded with high axial force and low moment. CFT columns also possess great strength and remarkable ductility because the steel tube provides a confinement to the concrete inside the tube, while the concrete inhibits local buckling of the steel tube. The construction techniques filling the concrete in tubes make CFT columns an economical alternative, compared to other composite column. A significant reduction in the steel weight using CFT columns compared to using bare steel columns was reported [1]. Corresponding author. Tel.: +886-2-2290-0312; fax: +886-2-22964145. E-mail address: [email protected] (T.-S. Moon).
0141-0296/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2004.06.016
The reduction is attributed to the concrete in steel tube increases the column stiffness. Japan has widely used internal, external, or through plate diaphragms in the connections [2]. The use of through plate diaphragms may be the most efficient form of strengthening CFT connections. The through plate diaphragm should have a hole in it to fill the concrete in tube column. The connections with internal or through diaphragms are commonly used for the highrise columns because the concrete is cast from bottom to top. The concrete is usually cast by dropping from top using a pumping pipe for the basement or low story columns. The internal diaphragms interrupt the pouring concrete from top and delay the working speed. The external diaphragm or T-stiffener types are more efficient for the manufacturing connections and casting concrete. The purpose of this paper is to investigate the behavior on CFT columns to H-beam connections stiffened with T-stiffeners. Various types of connections using CFT columns have been reported. Elremaily and Azizinamini
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Nomenclature Bb db Db Fyf hFys vFys Hs Ls PI PII PIII bM p
Mu tc tf tv th b hp hp Rhp hs g
width of beam flange distance between tension and compression couple force in a beam depth of column flange yield strength of beam flange. yield strength of horizontal stiffener yield strength of vertical stiffener height of vertical stiffener length of horizontal stiffener design strength by the equation of failure mode I design strength by the equation of failure mode II design strength by the equation of failure mode III beam plastic moment ultimate moment thickness of column tube thickness of beam flange thickness of vertical stiffener thickness of horizontal stiffener beam plastic rotation total plastic rotation cumulative plastic rotation maximum skeleton rotation normalized cumulative plastic rotation (=Rhp/hy)
conducted a study on CFT column to H-beam connections with HT bolts in rigid frames [3]. Masuda et al. and Kimura and Matsui investigated the experimental performance of CFT column to H-beam connections with vertical stiffener plates [4,5]. Lee et al. conducted a similar study, experimentally and analytically determining the effect of T-stiffeners on the behavior of boxcolumn to I-beam connections using the finite element method [6]. Likewise, Shin et al. studied the effect of T-stiffeners and bent plates on the behavior of boxcolumn to H-beam connections [7]. The specimens subjected to cyclic loading showed that failure was initiated by a welding crack and deformation capacity was insufficient. The pulling force generated in the vertical stiffeners caused a brittle failure on the corner of column because the vertical stiffeners were welded on the cold-formed round corner. In previous studies [8–10], the exterior face of vertical stiffeners was aligned with the centerline of the column wall and test results showed better hysteretic behavior. The previous welding details include the welding of vertical stiffener at the round corner of cold-formed square tube, which is dangerous because of large plastic strains at the cor-
ners due to cold forming. Therefore, in present study, instead of cold-formed tube column, built-up box column was used for effective transferring the force from the vertical stiffener to the column. The overall objective of this paper is to develop a connection design method using T-stiffeners for a CFT column to H-beam connection. Specific goal of this study are to (1) gain a better understanding of the inelastic behavior of the T-stiffeners; (2) develop the design equations for a connection with T-stiffeners; (3) investigate the effect of T-stiffeners in the force transfer mechanism around a connection; (4) assess the deformation capacity of the connections with different Tstiffener sizes. These objectives are addressed through the sub-assemblage connection tests and non-linear finite element analyses using ABAQUS [11].
2. T-stiffener design concept Most moment-resisting connections act both in shear and moment resistance simultaneously. In a rigid connection, the moment transferred from the beam to the column takes the form of a couple consisting of tensile and compressive forces in the beam flanges. In CFT column to H-beam connections with T-stiffeners, compressive force is resisted by both the stiffeners and the filled concrete. However, the tension force is resisted only by the T-stiffeners. A T-stiffener is composed of a vertical stiffener and a horizontal stiffener. Axial force in the beam flange transfers from the horizontal and vertical stiffeners to the webs of column because the column flange in tension is relatively flexible due to out-of-plane bending. The distribution of stresses and the intensity of stress concentration can be quite sensitive to the choice of T-stiffener details. Therefore, a simplified load path flow model must be formulated to derive a simple design procedure. For the T-stiffener connections, the following assumptions were made: (a) the moment developed at the connection (Fig. 1) should be equal to the plastic moment capacity bMp of a beam, and; (b) the moment is carried by the beam flanges in the form of a couple force Tp (Eq. (1)). Hence a force Tp/2 applied to a T-stiffener and should be carried by a horizontal and a vertical stiffener. The shear force V can be carried through the welding of web as assumed as a conventional design method. In calculating plastic moment capacity, the expected yield strength or strain hardening factors guided by AISC Seismic Provisions [12] were not included in this study. In Lee’s design methods [6], the shear capacity of the failure plane, Tp/2, was assumed as a summation of the shear capacity of stiffener web and the shear capacity at the column wall, which means that the length and the height of T-stiffener are dependent
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tensile failure of beam flange (Mode III). Tp ¼ b Mp =db
ð1Þ
Tp ¼ PI ¼ th Ls 0:6h Fys 2 Tp Mode II : ¼ PII ¼ tv Hs v Fys 2 Mode III : Tp ¼ PIII ¼ tf Bf Fyf
ð2Þ
Mode I :
ð3Þ ð4Þ
The main parameters of this study are the ratios of T-stiffener strength to beam strength as shown in Table 1. If the ratio of horizontal stiffener strength is smaller than 100%, the stiffener strength is not enough to resist beam strength, which may result in failure Mode I. Even though this ratio is larger than 100%, premature failure might occur due to the stress concentration at the welding. A test conducted by Moon [9], also revealed that welding failure between the horizontal stiffener and beam flange triggered the failure of connections. All specimens tested in this study were designed for the above-mentioned criteria for different failure modes.
Fig. 1.
Force transfer mechanism of connection with T-stiffener.
3. Experimental program 3.1. Test specimens
design parameters. Thus, the height of the vertical stiffener is calculated using the force derived by subtracting the shear capacity of the horizontal stiffener from the force Tp/2. However, previous study [9] demonstrated that the force Tp/2 would be carried as a serial type by a vertical and a horizontal stiffener, respectively. Simple design criteria are therefore suggested for connections with T-stiffeners. These design methods are based on the three failure modes of connections: (1) shear failure of horizontal stiffener (Mode I); (2) tensile failure of vertical stiffener at column web (Mode II); (3)
Six specimens were manufactured (Table 1). They were designed for the following test parameters: (a) the ratios of the horizontal stiffener strength to beam flange strength was 70% and 130%; (b) the ratios of the vertical stiffener strength to beam flange strength was 70%, 100% and 130%; and (c) two hot-rolled beams were H-588 ðbeam depthÞ 300 ðflange widthÞ 12 ðweb thicknessÞ 20 (flange thickness) and H-506 201 11 19. Figs. 1 and 2 show the details of typical specimen and the horizontal and vertical stiffener of the T-stiffener in which Hs and Ls are the height of the vertical stiffeners and the length of the horizontal stiffeners of
Table 1 Schedule of specimens Specimen
TS-1 TS-2 TS-3 TS-4 TS-5 TS-6 a b
Horizontal stiffener
Vertical stiffener
Strength ratio a (%)
Length (Ls) (mm)
Strength ratio b(%)
Height (Hs) (mm)
70 130 130 70 130 130
240 440 440 160 300 300
100 70 100 130 70 130
360 230 360 280 150 280
This is the ratio of the horizontal stiffener strength to beam flange strength. This is the ratio of the vertical stiffener strength to beam flange strength.
H-beam and built-up tube
Expected failure mode
H-588 300 12 20 &-500 500 12
Mode I Mode II Mode III Mode I Mode II Mode III
H-506 201 11 19 &-500 500 12
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Fig. 2.
T-stiffener connection detail.
the T-stiffener, respectively. The materials used in the test were KS SS400 (similar to A36) mild steel plates for the T-stiffeners and beams. The horizontal stiffeners were welded to the beam flange using double-V-groove welds, and to the column flange using complete joint penetration (CJP) single-bevel-groove welds. The vertical stiffeners were welded to the horizontal stiffeners using double-bevel-groove welds and to the column
corner using single-bevel-groove-welds. Test specimens were fabricated using a built-up section so that the centerline of the column web was aligned with that of vertical stiffener. Welding electrodes with a specified v minimum CVN toughness of 80 J at 20 C were used. The coupon test results are shown in Table 2. The average measured compressive strength of the concrete used to fill the tube was 26.46 MPa.
Table 2 Tensile coupon test results Specimen
Yield stress (Mpa)
Ultimate stress (Mpa)
Yield ratio
Elongation (%)
H-588 300 12 20 Beam flange (20 mm) Beam web (12 mm)
306 319
445 440
0.69 0.73
24 28
H-506 201 11 19 Beam flange (19 mm) Beam web (11 mm)
288 308
420 445
0.69 0.69
37 30
&-500 500 12 Tube plate (12 mm)
271
458
0.59
26
T-stiffener Horizontal stiffener plate (20 mm) Horizontal stiffener plate (19 mm) Vertical stiffener plate (12 mm)
369 285 289
571 462 518
0.65 0.62 0.56
26 34 26
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3.2. Test setup All beam to column connection tests were conducted on specimens fabricated as cantilevers attached to CFT column stubs. There was no floor slab on beam. The test setup was designed to simulate the boundary conditions of a beam to column connection sub-assembly in a moment-resisting frame under typical lateral loading. The column was assumed to be pin-supported at midstory points and the beam was loaded at an assumed pin at its midspan. Fig. 3 shows the test setup and the dimensions of typical specimens. The distance measured from the column centerline to the loading point of actuator was 3500 mm. Strain gauges were applied to the specimens in order to measure strain distributions and principal strain magnitudes and directions at critical points. Displacement transducers were used to measure the beam tip displacement and rotation contributions of different parts of the specimen. The cyclic loads, producing lateral seismic load effect, were applied vertically through an actuator mounted between the reaction frame and the beam end. The axial force for concrete filled column was not applied. The loading protocol prescribed a quasi-static cyclic displacement pattern defined in terms of beam end displacement (Fig. 4). The specimens were tested with rotation amplitudes of 1hy, 3hy, 5hy, etc. (hy is the elastic rotation corresponding to the beam plastic moment, bMp.) This loading history was applied to each specimen until failure occurred.
Fig. 3.
Test setup.
Fig. 4. Loading protocol.
4. Experimental results 4.1. Global behavior Moments versus rotation relationships obtained from the tests are presented in Fig. 5. The horizontal axes are the total beam rotations (hm), calculated as the tip displacements divided by the distance between the loading point and the centerline of column (3500 mm). The vertical axes show the connection moment (Mm) normalized with respect to the beam plastic moment bMp. Table 3 presents a summary of experimental results for the six tests. The ratio of the ultimate moment to the beam plastic moment obtained for this test ranged from 0.96 to 1.47. The total plastic rotation hp is obtained by subtracting the elastic deformations of beam, column and panel zone and bhp is the plastic rotation of the beam only. The cumulative plastic rotations (Rhp) and the failure modes are included in the table. TS-1 and TS-4 specimens that had a horizontal stiffener to beam flange strength ratio of 70% failed at the horizontal stiffeners as expected and showed poor hysteresis characteristics. During the three cycles, a welding crack between the beam flange and the horizontal stiffener appeared in specimen TS-1. When the crack fully opened, the lower beam flange fractured. In specimen TS-4, a welding crack appeared between the upper beam flange and the horizontal stiffener in the second cycle, as well as between the lower beam flange and the horizontal stiffener during the third cycle. The specimen underwent a large deformation in a ductile manner after loosing the full strength because out-ofplane bending of column flange could contribute to the load-carrying capacity. TS-2 had the poorest experimental performance of all the specimens. During the second cycle, TS-2 failed across the heat-affected zone of welding between the vertical stiffener and column web and lost load-carrying capacity. The maximum moment was slightly higher than the beam plastic moment. On the other hand, specimen TS-5 was ductile despite its vertical stiffener to beam flange strength ratio being 70%. It resulted in a stable hysteretic characteristic as shown in Fig. 5(e). The maximum
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Fig. 5. Normalized moment versus rotation relationships.
Table 3 Test results Specimen
Mu (kN m)
Mp (kN m)
Mu/bMp
bhp
TS-1 TS-2 TS-3 TS-4 TS-5 TS-6
1371 1623 2038 717 1076 1051
1388 1388 1388 746 746 746
(rad)
0.99 1.17 1.47 0.96 1.44 1.41
0.026 0.008 0.030 0.021 0.043 0.029
hp (rad)
Rhp (rad)
g
Failure mode
0.028 0.008 0.033 0.023 0.045 0.031
0.093 0.008 0.170 0.051 0.263 0.200
14 1 27 7 37 28
Horizontal stiffener failure Vertical stiffener failure Beam flange failure Horizontal stiffener failure Vertical stiffener failure Beam flange failure
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moment was about 1076 kN m, which was about 1.44 times the beam plastic moment. Local buckling of beam flange and web also occurred. It should be noted that the plastic modulus of TS-5 is much smaller than that of TS-2. It is therefore important to design the connection detail carefully for the larger beam size so that the deformation capacity increases in the connections. The TS-3 and TS-6 specimens where the horizontal to beam flange strength ratio was 130% carried loads well above the plastic moment of the beam. This indicates a ductile hysteresis loop. The maximum moment for TS-3 and TS-6 was 1.47 (2038 kN) and 1.41 (1051 KN) times the beam plastic moment, respectively. In terms of total plastic rotation, specimen TS-5 possessed the best rotation capacity (0.045 rad) and specimen TS-2 had the worst (0.008 rad). Three specimens, TS-3, TS-5 and TS-6, failed after reaching the total plastic rotation 0.03 rad. 4.2. Failure mode The failure modes of the connections included shear failure in the horizontal stiffener, tensile failure in the vertical stiffener, and beam flange failure and local buckling in beam. It is important to control the T-stiffener strength so that the hinge forms primarily in the beam. Typical load–displacement responses of the specimens depended on the sequence of failure of the vertical and horizontal stiffeners of the T-stiffeners. The specimens TS-1 and TS-2 failed by failure mode I as expected (Fig. 6(a)). Regardless of the height of the vertical stiffener, the strength of the connection was governed by the strength of horizontal stiffener. The cracks were detected at the end of welding between the horizontal stiffener and beam flange and then progressed inward along the welding line to the column flange. However, other specimens having horizontal stiffener to beam flange strength ratio of 130% did not experience out-of-plane deformation of the column flange. They also did not exhibit failure of the horizontal stiffener. On the other hand, failure mode II is a vertical stiffener failure of specimens having a vertical stiffener to beam flange strength ratio of 70%. As shown in Fig. 6(b) specimens TS-2 and TS-5 failed by the fracture of the welding or the heat-affected zone on the column web side. These test specimens were fabricated from built-up section, where the column web met the vertical stiffener directly. Thus, the moment from the beam directly transferred tensile force into the column web via the vertical stiffener. The hysteretic response of the TS-2 specimen stopped when vertical stiffener fractured abruptly. Fig. 6(c) shows failure mode III as a tensile and local buckling failure mode of the beam. This type of con-
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nection should be improved to avoid brittle failure at welding or the heat-affected zone. TS-3 and TS-6 specimens were reinforced with a vertical and horizontal stiffener to beam flange strength ratio of 100% or 130%, respectively. These specimens exhibited local buckling and yielding in the beam flange and the web near T-stiffener. Consequently, in mode III, the plastic hinge mechanism of the beam formed away from the face of the column. The yielding zone of the beam, located in the T-stiffener end, ensured the development of reliable hysteresis behavior and the dissipation of energy by the extensive yielding of the beam (Fig. 6(c)). But, specimen TS-3 ultimately failed during the fourth cycle by the fracture of beam flange. It was likely that this type of fracture started due to the stress concentration at the end of horizontal stiffener. The horizontal stiffener plays a more effective role than the vertical stiffener in increasing the strength of connections because the shear stress in the horizontal stiffener changes drastically. Abrupt changes in beam flange area are not permitted in plastic hinge regions for the special moment frames [12]. Further study such as tapered horizontal stiffener is needed to reduce the stress concentrations and develop more stable plastic hinges. 4.3. Assessment of ductility Moment–rotation response of the specimens in this test depended on the critical elements such as the vertical stiffener in the tension zone, horizontal stiffener, and beam flange. The response of all specimens except TS-2 was fair before fracture occurred, as evidenced by the shape of hysteresis loops in Fig. 5. Table 3 shows that the total plastic rotational capacity of all specimens ranged between 0.008 and 0.045 rad. Fig. 7 shows the total plastic rotation and beam plastic rotation when failure first occurs either at the vertical stiffener, the horizontal stiffener, or beam flange. Since the panel zone was designed to be very strong, its plastic rotation was zero and consequently not shown in this paper. It should be noted that the plastic rotational capacities of specimens TS-3 and TS-6 showed 0.033 and 0.032 rad, respectively. As shown in Fig. 7, for all specimens, a beam plastic rotation was about 90% or higher than 90% comparing the total plastic rotation. This indicated that the deformation of the beam to column connections is mainly provided by the beam deformation, supposing that the plastic hinge will be capable of forming away from the column face. The concept of a skeleton curve has been commonly adopted in characterizing the deformation capacity of steel members subjected to load reversal in Japan [13]. The skeleton curve constructed from moment versus rotation hysteresis loops can also be defined here. Hysteresis loops were assumed to increase their rotation
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Fig. 6.
Failure modes after test :(a) failure mode I; (b) failure mode II; (c) failure mode III.
range incrementally with each cycle (Fig. 8). When the load during the second cycle exceeds the peak load on the first cycle, the exceeding portion of the loop during the second cycle is connected to the point at the peak load of the first cycle. The same process is repeated until the highest moment is reached. In this procedure, a skeleton curve was defined and constructed for each of the positive and negative loading directions. The skeleton rotation is also useful when comparing the ductility capacity and energy dissipation capacity of hysteresis curves of various types of specimens. Fig. 9
plots the maximum skeleton rotation (hs) and the corresponding cumulative plastic rotation (Rhp) for all test specimens [14]. Fig. 9 also shows the larger of the posi tive (hþ s ) and negative (hs ) maximum skeleton rotations. In general, a larger cumulative plastic rotation gave a large maximum skeleton rotation. The maximum skeleton rotation and cumulative plastic rotation were relatively larger for specimens reinforced with T-stiffeners having 130% of strength to beam flange.
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Fig. 7.
Plastic rotation.
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the transmission of force between the surfaces occurs when they are in contact. Material non-linearity is accounted for in the monotonic analyses through an isotropic classical plasticity model, based on the Von Mises yield criterion and associated plastic flow. Isotropic hardening was assumed for monotonic analyses, whereas kinematic hardening was assumed for cyclic analyses. Isotropic hardening rule mechanism that produces hardening behavior in tension was assumed to act equally in compression and contradicts the Bauschinger effect observed in experiments (Fig. 10(a)). Fig. 10(b) shows the kinematic hardening rule, i.e., the elastic region moves around in stress space as a rigid body depending on aa’ line and exhibits the idealized Bauschinger effect [15]. Therefore, the kinematic hardening rule considered non-linearity, as such the Bauschinger effect was adopted in this study. Fig. 11 shows a global finite element model. 5.2. Verification
Fig. 8.
Method for drawing the skeleton curve.
Fig. 12 shows a comparison between the experimental and analytical results for specimen TS-3 and TS-5. A close examination of the figures reveals that the analytical hysteretic loops had slightly sharper corners than the experimentally observed ones, although the ultimate load and initial stiffness are well repre-
Fig. 9. Maximum skeleton rotation versus cumulative plastic rotation relationships.
5. Non-linear finite element analysis 5.1. Element modeling The analyses were conducted using the computer program ABAQUS [11]. The finite element model used in this study was an eight-node brick (C3D8R) reduced integration element. In this study, contact surfaces were used in consideration of the filling effect of the concrete. The mechanical interaction of an interface between surfaces involves contact and separation, i.e.,
Fig. 10. Hardening rule: (a) isotropic hardening; (b) kinematic hardening.
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Fig. 11. Detail of finite element model.
Fig. 12. Comparison of test and finite element analysis: (a) TS-3; (b) TS-5.
sented. This is possibly because residual stresses were not included in the analysis. Fig. 13 shows the Von Mises stress contours at 0.03 rad beam plastic rotation (BPR) for TS-1, TS-3, and TS-5. Stress concentration occurred at the horizontal stiffener for TS-1 (Fig. 13(a)), vertical stiffener for TS-5 (Fig. 13(b)), and beam flange for TS-3 (Fig. 13(c)). These stress concentrations in the analytical results showed reasonably good correlation with the failure modes from the test results. According to the Fig. 13, the proposed strength Eqs. (2)–(4) based on the failure mode can predict accurately the strength of the connections reinforced with T-stiffeners. These observations clearly indicate that T-stiffener designed properly for the proposed strength equations leads to the formation of the plastic hinge in the beam section away from the column face. The energy dissipated within the beam
Fig. 13. Von Mises stress contour at 0.03 BPR: (a) TS-1; (b) TS-5; (c) TS-3.
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section is deemed to be more reliable than that dissipated starting from the stiffeners.
6. Conclusions Based on both experimental results and analytical studies, the following conclusions can be drawn concerning moment connections reinforced by the T-stiffener. The failure modes of the connections included three modes of the shear failure at the horizontal stiffener, the tensile failure at the vertical stiffener, and the tensile failure and buckling at beam. For test and analytical results, the proposed strength equations based on the failure mode can predict the failure modes of the connections reinforced with T-stiffeners. In T-stiffener connections, while the vertical stiffener only effectively transfers load introduced from beam to column, the horizontal stiffener plays a more critical role in increasing the ultimate moment and deformation capacity. The mean plastic rotation of TS-3 and TS-6 was 0.032 rad. The average plastic rotation exceeded 0.03 rad, required in special moment frame (SMF), even though this test was not based on a specified loading protocol [12]. The strain hardening factor and the expected yield strength factor (Ry) in designing T-stiffeners need to be considered to obtain a plastic hinge in beam [12]. The test results indicate that specimens with the strength ratio of 130% for the proposed strength equations are suitable for developing the plastic hinge at the beam and enhancing the deformation capacity of the connections.
Acknowledgements This study is part of the research on Concrete-Filled Column Connections being carried out in the STructure RESearch Station (STRESS) and POSCO. The author gratefully acknowledges the support of STRESS and POSCO.
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References [1] Zhang H, Ricles J. An analytical study of seismic resistant CFT MRFs. Research Report, Department of Civil Environmental Engineering, Lehigh University, Bethlehem, PA; 1996. [2] AIJ standard for structural calculation of steel reinforced concrete structures: allowable stress design and horizontal load-carrying capacity. Revised 2001, AIJ 2001. [3] Elremaily A, Azizinamini A. Experimental behavior of steel beam to CFT column connections. Journal of Constructional Steel Research 2001;57(10):1099–119. [4] Masuda H, Tanaka A, Ishimura R, Hirai K, Sasaji S. Experimental study on the statical characteristics of perimeter beam-toSHS column connections using vertical stiffener plates. Journal of structural and Construction Engineering, AIJ 2000;534(8): 167–74. [5] Kimura J, Matsui C. Structural performance of H-shaped steel beam to square tube steel column connection using vertical stiffeners. Summaries of technical papers of Annual Meeting, AIJ 2000. p. 631–32. [6] Lee SL, Ting LC, Shanmugam NE. Use of external T-stiffeners in box-column to beam connections. Journal of Construction Steel Research 1993;26:77–98. [7] Shin KJ, Oh YS, Moon TS. Test of concrete-filled box column to H-beam connections. In: Proceedings of Pacific Structural Steel Conference (PSSC). Seoul (South Korea): Korean Society of Steel Construction (KSSC); 1998, p. 881–6. [8] Oh YS. CFT column to H-beam connections with T-stiffeners. In: Report. No. STRESS-98-07, Seoul (South Korea): Advanced Structure Research Station (STRESS), Hanyang University; 1999. [9] Moon TS, Oh YS, Shin KJ, Kang CH. Hystersis behavior of CFT column to H-beam connections with external T-stiffeners and penetrated elements. Engineering Structure 2001;23(9): 1194–201. [10] Kang CH, Chae YS, Shin KJ, Oh YS, Moon TS. CFT columnto-beam flange connections with external stiffeners. Second International Symposium on Steel Structures 21–22 November, Seoul, South Korea. 2002, p. 123–30. [11] ABAQUS standard user’s manual-version 5.8. Pawtucket (RI): Hibbit, Karlsson, and Sorensen; 1998. [12] Seismic provisions for structural steel buildings. 1st ed. AISC, Chicago; 1997. [13] Kato B, Akiyama H. The ultimate strength of the steel beam-tocolumn (part 4). Transactions of AIJ 1968;151:15–20. [14] Nakashima M, Suita K, Morisako K. Test of welded beam–column subassemblies. II: detail behavior. Journal of Structural Engineering, ASCE 1998;124(11):1245–52. [15] Chen WF, Han DJ. Plasticity for structural engineers. SpringerVerlag; 1998.
Behavior of welded CFT column to H-beam connections with external stiffeners Kyung-Jae Shin a, Young-Ju Kim b, Young-Suk Oh c, Tae-Sup Moon b, a
b
Department of Architectural Engineering, Hannam University, 133 Ojung-dong taeduk-gu, Taejon 306-791, South Korea Department of Architectural Engineering, Hanyang University, 17 Haengdang-dong, Seongdong-gu, Seoul 133-791, South Korea c Department of Architectural Engineering, Taejon University, 96-3 Yongun-dong, Tong-gu, Taejon 300-716, South Korea Received 5 September 2003; received in revised form 21 June 2004; accepted 23 June 2004
Abstract This paper focuses on the experimental and analytical behavior of concrete filled tubular (CFT) column to H-beam welded moment connections with external T-stiffeners. Six full-scale specimens were tested cyclically. All specimens were designed and manufactured using the following main test parameters: (a) the strength ratio of the horizontal stiffener to the beam flange; and (b) the strength ratio of the vertical stiffener to the beam flange. Three types of failure modes, horizontal stiffeners failure, vertical stiffener failure and beam failure, were obtained. The connections reinforced with T-stiffeners having 130% of strength to beam flanges showed stable hysteretic behavior and good ductility. The ABAQUS finite element package was used to simulate the experimental behavior. The results obtained from the finite element were evaluated by comparing the load–displacement responses and the potential of failure modes. # 2004 Elsevier Ltd. All rights reserved. Keywords: Concrete filled tube (CFT); T-stiffener; Moment-resisting frame; Plastic rotation
1. Introduction Steel–concrete composite materials have been widely used in framing systems for high-rise buildings. One particular type of composite construction is concrete filled steel tube (CFT). CFT columns have been found to have advantages especially for the column loaded with high axial force and low moment. CFT columns also possess great strength and remarkable ductility because the steel tube provides a confinement to the concrete inside the tube, while the concrete inhibits local buckling of the steel tube. The construction techniques filling the concrete in tubes make CFT columns an economical alternative, compared to other composite column. A significant reduction in the steel weight using CFT columns compared to using bare steel columns was reported [1]. Corresponding author. Tel.: +886-2-2290-0312; fax: +886-2-22964145. E-mail address: [email protected] (T.-S. Moon).
0141-0296/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2004.06.016
The reduction is attributed to the concrete in steel tube increases the column stiffness. Japan has widely used internal, external, or through plate diaphragms in the connections [2]. The use of through plate diaphragms may be the most efficient form of strengthening CFT connections. The through plate diaphragm should have a hole in it to fill the concrete in tube column. The connections with internal or through diaphragms are commonly used for the highrise columns because the concrete is cast from bottom to top. The concrete is usually cast by dropping from top using a pumping pipe for the basement or low story columns. The internal diaphragms interrupt the pouring concrete from top and delay the working speed. The external diaphragm or T-stiffener types are more efficient for the manufacturing connections and casting concrete. The purpose of this paper is to investigate the behavior on CFT columns to H-beam connections stiffened with T-stiffeners. Various types of connections using CFT columns have been reported. Elremaily and Azizinamini
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Nomenclature Bb db Db Fyf hFys vFys Hs Ls PI PII PIII bM p
Mu tc tf tv th b hp hp Rhp hs g
width of beam flange distance between tension and compression couple force in a beam depth of column flange yield strength of beam flange. yield strength of horizontal stiffener yield strength of vertical stiffener height of vertical stiffener length of horizontal stiffener design strength by the equation of failure mode I design strength by the equation of failure mode II design strength by the equation of failure mode III beam plastic moment ultimate moment thickness of column tube thickness of beam flange thickness of vertical stiffener thickness of horizontal stiffener beam plastic rotation total plastic rotation cumulative plastic rotation maximum skeleton rotation normalized cumulative plastic rotation (=Rhp/hy)
conducted a study on CFT column to H-beam connections with HT bolts in rigid frames [3]. Masuda et al. and Kimura and Matsui investigated the experimental performance of CFT column to H-beam connections with vertical stiffener plates [4,5]. Lee et al. conducted a similar study, experimentally and analytically determining the effect of T-stiffeners on the behavior of boxcolumn to I-beam connections using the finite element method [6]. Likewise, Shin et al. studied the effect of T-stiffeners and bent plates on the behavior of boxcolumn to H-beam connections [7]. The specimens subjected to cyclic loading showed that failure was initiated by a welding crack and deformation capacity was insufficient. The pulling force generated in the vertical stiffeners caused a brittle failure on the corner of column because the vertical stiffeners were welded on the cold-formed round corner. In previous studies [8–10], the exterior face of vertical stiffeners was aligned with the centerline of the column wall and test results showed better hysteretic behavior. The previous welding details include the welding of vertical stiffener at the round corner of cold-formed square tube, which is dangerous because of large plastic strains at the cor-
ners due to cold forming. Therefore, in present study, instead of cold-formed tube column, built-up box column was used for effective transferring the force from the vertical stiffener to the column. The overall objective of this paper is to develop a connection design method using T-stiffeners for a CFT column to H-beam connection. Specific goal of this study are to (1) gain a better understanding of the inelastic behavior of the T-stiffeners; (2) develop the design equations for a connection with T-stiffeners; (3) investigate the effect of T-stiffeners in the force transfer mechanism around a connection; (4) assess the deformation capacity of the connections with different Tstiffener sizes. These objectives are addressed through the sub-assemblage connection tests and non-linear finite element analyses using ABAQUS [11].
2. T-stiffener design concept Most moment-resisting connections act both in shear and moment resistance simultaneously. In a rigid connection, the moment transferred from the beam to the column takes the form of a couple consisting of tensile and compressive forces in the beam flanges. In CFT column to H-beam connections with T-stiffeners, compressive force is resisted by both the stiffeners and the filled concrete. However, the tension force is resisted only by the T-stiffeners. A T-stiffener is composed of a vertical stiffener and a horizontal stiffener. Axial force in the beam flange transfers from the horizontal and vertical stiffeners to the webs of column because the column flange in tension is relatively flexible due to out-of-plane bending. The distribution of stresses and the intensity of stress concentration can be quite sensitive to the choice of T-stiffener details. Therefore, a simplified load path flow model must be formulated to derive a simple design procedure. For the T-stiffener connections, the following assumptions were made: (a) the moment developed at the connection (Fig. 1) should be equal to the plastic moment capacity bMp of a beam, and; (b) the moment is carried by the beam flanges in the form of a couple force Tp (Eq. (1)). Hence a force Tp/2 applied to a T-stiffener and should be carried by a horizontal and a vertical stiffener. The shear force V can be carried through the welding of web as assumed as a conventional design method. In calculating plastic moment capacity, the expected yield strength or strain hardening factors guided by AISC Seismic Provisions [12] were not included in this study. In Lee’s design methods [6], the shear capacity of the failure plane, Tp/2, was assumed as a summation of the shear capacity of stiffener web and the shear capacity at the column wall, which means that the length and the height of T-stiffener are dependent
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tensile failure of beam flange (Mode III). Tp ¼ b Mp =db
ð1Þ
Tp ¼ PI ¼ th Ls 0:6h Fys 2 Tp Mode II : ¼ PII ¼ tv Hs v Fys 2 Mode III : Tp ¼ PIII ¼ tf Bf Fyf
ð2Þ
Mode I :
ð3Þ ð4Þ
The main parameters of this study are the ratios of T-stiffener strength to beam strength as shown in Table 1. If the ratio of horizontal stiffener strength is smaller than 100%, the stiffener strength is not enough to resist beam strength, which may result in failure Mode I. Even though this ratio is larger than 100%, premature failure might occur due to the stress concentration at the welding. A test conducted by Moon [9], also revealed that welding failure between the horizontal stiffener and beam flange triggered the failure of connections. All specimens tested in this study were designed for the above-mentioned criteria for different failure modes.
Fig. 1.
Force transfer mechanism of connection with T-stiffener.
3. Experimental program 3.1. Test specimens
design parameters. Thus, the height of the vertical stiffener is calculated using the force derived by subtracting the shear capacity of the horizontal stiffener from the force Tp/2. However, previous study [9] demonstrated that the force Tp/2 would be carried as a serial type by a vertical and a horizontal stiffener, respectively. Simple design criteria are therefore suggested for connections with T-stiffeners. These design methods are based on the three failure modes of connections: (1) shear failure of horizontal stiffener (Mode I); (2) tensile failure of vertical stiffener at column web (Mode II); (3)
Six specimens were manufactured (Table 1). They were designed for the following test parameters: (a) the ratios of the horizontal stiffener strength to beam flange strength was 70% and 130%; (b) the ratios of the vertical stiffener strength to beam flange strength was 70%, 100% and 130%; and (c) two hot-rolled beams were H-588 ðbeam depthÞ 300 ðflange widthÞ 12 ðweb thicknessÞ 20 (flange thickness) and H-506 201 11 19. Figs. 1 and 2 show the details of typical specimen and the horizontal and vertical stiffener of the T-stiffener in which Hs and Ls are the height of the vertical stiffeners and the length of the horizontal stiffeners of
Table 1 Schedule of specimens Specimen
TS-1 TS-2 TS-3 TS-4 TS-5 TS-6 a b
Horizontal stiffener
Vertical stiffener
Strength ratio a (%)
Length (Ls) (mm)
Strength ratio b(%)
Height (Hs) (mm)
70 130 130 70 130 130
240 440 440 160 300 300
100 70 100 130 70 130
360 230 360 280 150 280
This is the ratio of the horizontal stiffener strength to beam flange strength. This is the ratio of the vertical stiffener strength to beam flange strength.
H-beam and built-up tube
Expected failure mode
H-588 300 12 20 &-500 500 12
Mode I Mode II Mode III Mode I Mode II Mode III
H-506 201 11 19 &-500 500 12
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Fig. 2.
T-stiffener connection detail.
the T-stiffener, respectively. The materials used in the test were KS SS400 (similar to A36) mild steel plates for the T-stiffeners and beams. The horizontal stiffeners were welded to the beam flange using double-V-groove welds, and to the column flange using complete joint penetration (CJP) single-bevel-groove welds. The vertical stiffeners were welded to the horizontal stiffeners using double-bevel-groove welds and to the column
corner using single-bevel-groove-welds. Test specimens were fabricated using a built-up section so that the centerline of the column web was aligned with that of vertical stiffener. Welding electrodes with a specified v minimum CVN toughness of 80 J at 20 C were used. The coupon test results are shown in Table 2. The average measured compressive strength of the concrete used to fill the tube was 26.46 MPa.
Table 2 Tensile coupon test results Specimen
Yield stress (Mpa)
Ultimate stress (Mpa)
Yield ratio
Elongation (%)
H-588 300 12 20 Beam flange (20 mm) Beam web (12 mm)
306 319
445 440
0.69 0.73
24 28
H-506 201 11 19 Beam flange (19 mm) Beam web (11 mm)
288 308
420 445
0.69 0.69
37 30
&-500 500 12 Tube plate (12 mm)
271
458
0.59
26
T-stiffener Horizontal stiffener plate (20 mm) Horizontal stiffener plate (19 mm) Vertical stiffener plate (12 mm)
369 285 289
571 462 518
0.65 0.62 0.56
26 34 26
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3.2. Test setup All beam to column connection tests were conducted on specimens fabricated as cantilevers attached to CFT column stubs. There was no floor slab on beam. The test setup was designed to simulate the boundary conditions of a beam to column connection sub-assembly in a moment-resisting frame under typical lateral loading. The column was assumed to be pin-supported at midstory points and the beam was loaded at an assumed pin at its midspan. Fig. 3 shows the test setup and the dimensions of typical specimens. The distance measured from the column centerline to the loading point of actuator was 3500 mm. Strain gauges were applied to the specimens in order to measure strain distributions and principal strain magnitudes and directions at critical points. Displacement transducers were used to measure the beam tip displacement and rotation contributions of different parts of the specimen. The cyclic loads, producing lateral seismic load effect, were applied vertically through an actuator mounted between the reaction frame and the beam end. The axial force for concrete filled column was not applied. The loading protocol prescribed a quasi-static cyclic displacement pattern defined in terms of beam end displacement (Fig. 4). The specimens were tested with rotation amplitudes of 1hy, 3hy, 5hy, etc. (hy is the elastic rotation corresponding to the beam plastic moment, bMp.) This loading history was applied to each specimen until failure occurred.
Fig. 3.
Test setup.
Fig. 4. Loading protocol.
4. Experimental results 4.1. Global behavior Moments versus rotation relationships obtained from the tests are presented in Fig. 5. The horizontal axes are the total beam rotations (hm), calculated as the tip displacements divided by the distance between the loading point and the centerline of column (3500 mm). The vertical axes show the connection moment (Mm) normalized with respect to the beam plastic moment bMp. Table 3 presents a summary of experimental results for the six tests. The ratio of the ultimate moment to the beam plastic moment obtained for this test ranged from 0.96 to 1.47. The total plastic rotation hp is obtained by subtracting the elastic deformations of beam, column and panel zone and bhp is the plastic rotation of the beam only. The cumulative plastic rotations (Rhp) and the failure modes are included in the table. TS-1 and TS-4 specimens that had a horizontal stiffener to beam flange strength ratio of 70% failed at the horizontal stiffeners as expected and showed poor hysteresis characteristics. During the three cycles, a welding crack between the beam flange and the horizontal stiffener appeared in specimen TS-1. When the crack fully opened, the lower beam flange fractured. In specimen TS-4, a welding crack appeared between the upper beam flange and the horizontal stiffener in the second cycle, as well as between the lower beam flange and the horizontal stiffener during the third cycle. The specimen underwent a large deformation in a ductile manner after loosing the full strength because out-ofplane bending of column flange could contribute to the load-carrying capacity. TS-2 had the poorest experimental performance of all the specimens. During the second cycle, TS-2 failed across the heat-affected zone of welding between the vertical stiffener and column web and lost load-carrying capacity. The maximum moment was slightly higher than the beam plastic moment. On the other hand, specimen TS-5 was ductile despite its vertical stiffener to beam flange strength ratio being 70%. It resulted in a stable hysteretic characteristic as shown in Fig. 5(e). The maximum
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Fig. 5. Normalized moment versus rotation relationships.
Table 3 Test results Specimen
Mu (kN m)
Mp (kN m)
Mu/bMp
bhp
TS-1 TS-2 TS-3 TS-4 TS-5 TS-6
1371 1623 2038 717 1076 1051
1388 1388 1388 746 746 746
(rad)
0.99 1.17 1.47 0.96 1.44 1.41
0.026 0.008 0.030 0.021 0.043 0.029
hp (rad)
Rhp (rad)
g
Failure mode
0.028 0.008 0.033 0.023 0.045 0.031
0.093 0.008 0.170 0.051 0.263 0.200
14 1 27 7 37 28
Horizontal stiffener failure Vertical stiffener failure Beam flange failure Horizontal stiffener failure Vertical stiffener failure Beam flange failure
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moment was about 1076 kN m, which was about 1.44 times the beam plastic moment. Local buckling of beam flange and web also occurred. It should be noted that the plastic modulus of TS-5 is much smaller than that of TS-2. It is therefore important to design the connection detail carefully for the larger beam size so that the deformation capacity increases in the connections. The TS-3 and TS-6 specimens where the horizontal to beam flange strength ratio was 130% carried loads well above the plastic moment of the beam. This indicates a ductile hysteresis loop. The maximum moment for TS-3 and TS-6 was 1.47 (2038 kN) and 1.41 (1051 KN) times the beam plastic moment, respectively. In terms of total plastic rotation, specimen TS-5 possessed the best rotation capacity (0.045 rad) and specimen TS-2 had the worst (0.008 rad). Three specimens, TS-3, TS-5 and TS-6, failed after reaching the total plastic rotation 0.03 rad. 4.2. Failure mode The failure modes of the connections included shear failure in the horizontal stiffener, tensile failure in the vertical stiffener, and beam flange failure and local buckling in beam. It is important to control the T-stiffener strength so that the hinge forms primarily in the beam. Typical load–displacement responses of the specimens depended on the sequence of failure of the vertical and horizontal stiffeners of the T-stiffeners. The specimens TS-1 and TS-2 failed by failure mode I as expected (Fig. 6(a)). Regardless of the height of the vertical stiffener, the strength of the connection was governed by the strength of horizontal stiffener. The cracks were detected at the end of welding between the horizontal stiffener and beam flange and then progressed inward along the welding line to the column flange. However, other specimens having horizontal stiffener to beam flange strength ratio of 130% did not experience out-of-plane deformation of the column flange. They also did not exhibit failure of the horizontal stiffener. On the other hand, failure mode II is a vertical stiffener failure of specimens having a vertical stiffener to beam flange strength ratio of 70%. As shown in Fig. 6(b) specimens TS-2 and TS-5 failed by the fracture of the welding or the heat-affected zone on the column web side. These test specimens were fabricated from built-up section, where the column web met the vertical stiffener directly. Thus, the moment from the beam directly transferred tensile force into the column web via the vertical stiffener. The hysteretic response of the TS-2 specimen stopped when vertical stiffener fractured abruptly. Fig. 6(c) shows failure mode III as a tensile and local buckling failure mode of the beam. This type of con-
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nection should be improved to avoid brittle failure at welding or the heat-affected zone. TS-3 and TS-6 specimens were reinforced with a vertical and horizontal stiffener to beam flange strength ratio of 100% or 130%, respectively. These specimens exhibited local buckling and yielding in the beam flange and the web near T-stiffener. Consequently, in mode III, the plastic hinge mechanism of the beam formed away from the face of the column. The yielding zone of the beam, located in the T-stiffener end, ensured the development of reliable hysteresis behavior and the dissipation of energy by the extensive yielding of the beam (Fig. 6(c)). But, specimen TS-3 ultimately failed during the fourth cycle by the fracture of beam flange. It was likely that this type of fracture started due to the stress concentration at the end of horizontal stiffener. The horizontal stiffener plays a more effective role than the vertical stiffener in increasing the strength of connections because the shear stress in the horizontal stiffener changes drastically. Abrupt changes in beam flange area are not permitted in plastic hinge regions for the special moment frames [12]. Further study such as tapered horizontal stiffener is needed to reduce the stress concentrations and develop more stable plastic hinges. 4.3. Assessment of ductility Moment–rotation response of the specimens in this test depended on the critical elements such as the vertical stiffener in the tension zone, horizontal stiffener, and beam flange. The response of all specimens except TS-2 was fair before fracture occurred, as evidenced by the shape of hysteresis loops in Fig. 5. Table 3 shows that the total plastic rotational capacity of all specimens ranged between 0.008 and 0.045 rad. Fig. 7 shows the total plastic rotation and beam plastic rotation when failure first occurs either at the vertical stiffener, the horizontal stiffener, or beam flange. Since the panel zone was designed to be very strong, its plastic rotation was zero and consequently not shown in this paper. It should be noted that the plastic rotational capacities of specimens TS-3 and TS-6 showed 0.033 and 0.032 rad, respectively. As shown in Fig. 7, for all specimens, a beam plastic rotation was about 90% or higher than 90% comparing the total plastic rotation. This indicated that the deformation of the beam to column connections is mainly provided by the beam deformation, supposing that the plastic hinge will be capable of forming away from the column face. The concept of a skeleton curve has been commonly adopted in characterizing the deformation capacity of steel members subjected to load reversal in Japan [13]. The skeleton curve constructed from moment versus rotation hysteresis loops can also be defined here. Hysteresis loops were assumed to increase their rotation
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Fig. 6.
Failure modes after test :(a) failure mode I; (b) failure mode II; (c) failure mode III.
range incrementally with each cycle (Fig. 8). When the load during the second cycle exceeds the peak load on the first cycle, the exceeding portion of the loop during the second cycle is connected to the point at the peak load of the first cycle. The same process is repeated until the highest moment is reached. In this procedure, a skeleton curve was defined and constructed for each of the positive and negative loading directions. The skeleton rotation is also useful when comparing the ductility capacity and energy dissipation capacity of hysteresis curves of various types of specimens. Fig. 9
plots the maximum skeleton rotation (hs) and the corresponding cumulative plastic rotation (Rhp) for all test specimens [14]. Fig. 9 also shows the larger of the posi tive (hþ s ) and negative (hs ) maximum skeleton rotations. In general, a larger cumulative plastic rotation gave a large maximum skeleton rotation. The maximum skeleton rotation and cumulative plastic rotation were relatively larger for specimens reinforced with T-stiffeners having 130% of strength to beam flange.
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Fig. 7.
Plastic rotation.
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the transmission of force between the surfaces occurs when they are in contact. Material non-linearity is accounted for in the monotonic analyses through an isotropic classical plasticity model, based on the Von Mises yield criterion and associated plastic flow. Isotropic hardening was assumed for monotonic analyses, whereas kinematic hardening was assumed for cyclic analyses. Isotropic hardening rule mechanism that produces hardening behavior in tension was assumed to act equally in compression and contradicts the Bauschinger effect observed in experiments (Fig. 10(a)). Fig. 10(b) shows the kinematic hardening rule, i.e., the elastic region moves around in stress space as a rigid body depending on aa’ line and exhibits the idealized Bauschinger effect [15]. Therefore, the kinematic hardening rule considered non-linearity, as such the Bauschinger effect was adopted in this study. Fig. 11 shows a global finite element model. 5.2. Verification
Fig. 8.
Method for drawing the skeleton curve.
Fig. 12 shows a comparison between the experimental and analytical results for specimen TS-3 and TS-5. A close examination of the figures reveals that the analytical hysteretic loops had slightly sharper corners than the experimentally observed ones, although the ultimate load and initial stiffness are well repre-
Fig. 9. Maximum skeleton rotation versus cumulative plastic rotation relationships.
5. Non-linear finite element analysis 5.1. Element modeling The analyses were conducted using the computer program ABAQUS [11]. The finite element model used in this study was an eight-node brick (C3D8R) reduced integration element. In this study, contact surfaces were used in consideration of the filling effect of the concrete. The mechanical interaction of an interface between surfaces involves contact and separation, i.e.,
Fig. 10. Hardening rule: (a) isotropic hardening; (b) kinematic hardening.
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Fig. 11. Detail of finite element model.
Fig. 12. Comparison of test and finite element analysis: (a) TS-3; (b) TS-5.
sented. This is possibly because residual stresses were not included in the analysis. Fig. 13 shows the Von Mises stress contours at 0.03 rad beam plastic rotation (BPR) for TS-1, TS-3, and TS-5. Stress concentration occurred at the horizontal stiffener for TS-1 (Fig. 13(a)), vertical stiffener for TS-5 (Fig. 13(b)), and beam flange for TS-3 (Fig. 13(c)). These stress concentrations in the analytical results showed reasonably good correlation with the failure modes from the test results. According to the Fig. 13, the proposed strength Eqs. (2)–(4) based on the failure mode can predict accurately the strength of the connections reinforced with T-stiffeners. These observations clearly indicate that T-stiffener designed properly for the proposed strength equations leads to the formation of the plastic hinge in the beam section away from the column face. The energy dissipated within the beam
Fig. 13. Von Mises stress contour at 0.03 BPR: (a) TS-1; (b) TS-5; (c) TS-3.
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section is deemed to be more reliable than that dissipated starting from the stiffeners.
6. Conclusions Based on both experimental results and analytical studies, the following conclusions can be drawn concerning moment connections reinforced by the T-stiffener. The failure modes of the connections included three modes of the shear failure at the horizontal stiffener, the tensile failure at the vertical stiffener, and the tensile failure and buckling at beam. For test and analytical results, the proposed strength equations based on the failure mode can predict the failure modes of the connections reinforced with T-stiffeners. In T-stiffener connections, while the vertical stiffener only effectively transfers load introduced from beam to column, the horizontal stiffener plays a more critical role in increasing the ultimate moment and deformation capacity. The mean plastic rotation of TS-3 and TS-6 was 0.032 rad. The average plastic rotation exceeded 0.03 rad, required in special moment frame (SMF), even though this test was not based on a specified loading protocol [12]. The strain hardening factor and the expected yield strength factor (Ry) in designing T-stiffeners need to be considered to obtain a plastic hinge in beam [12]. The test results indicate that specimens with the strength ratio of 130% for the proposed strength equations are suitable for developing the plastic hinge at the beam and enhancing the deformation capacity of the connections.
Acknowledgements This study is part of the research on Concrete-Filled Column Connections being carried out in the STructure RESearch Station (STRESS) and POSCO. The author gratefully acknowledges the support of STRESS and POSCO.
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