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Seismic behavior and retrofit of steel moment connections considering slab effects
Engineering Structures 26 (2004) 1993–2005 www.elsevier.com/locate/engstruct
Seismic behavior and retrofit of steel moment connections considering slab effects Young-Ju Kim a,, Sang-Hoon Oh b, Tae-Sup Moon a a b
Department of Architectural engineering, Hanyang University,17, Haengdang-dong, Seongdong-gu, Seoul,133-791, South Korea Research Institute of Industrial Science & Technology, 79, Youngcheon Dongtan, Hwasung, Kyoungkido, 445-810, South Korea Received 29 January 2004; received in revised form 19 July 2004; accepted 27 July 2004
Abstract A series of five full-scale subassemblages were tested to investigate the contribution of the slabs and the effects of three types of retrofit methods, no weld access hole, horizontal stiffener, and cover plate. The test included one bare steel specimen and four composite specimens with floor slab. The results that emphasize the influence of the composite slab on connection behavior and specific comments on the retrofit scheme are presented. The test result indicated that the strains near the bottom flange of the composite beam connections were several times larger than those of the bare steel beam connections, resulting in a higher potential of fracture. Therefore, the slab effects are detrimental to the seismic behavior of the connection and should be considered in the design. Horizontal stiffener detail of three retrofit schemes demonstrated very good potential in improving the ductility of composite connections in existing buildings. # 2004 Elsevier Ltd. All rights reserved. Keywords: Slab effect; Retrofit; Deformation capacity; Strain concentration; Weld access hole; Horizontal stiffener
1. Introduction During the Northridge earthquake in California in 1994 and the Kobe earthquake in 1995, many steel buildings suffered from fracturing in the welded moment connections [1,2]. The performance of steel buildings during these earthquakes has raised questions regarding the reliability of the current connection method. The current connection design may be susceptible to failure in an earthquake. Investigation of these fractured connections indicated that most of the fractures occurred at the bottom flange. After the earthquake, many different types of connection methods have been suggested to solve the problem both in Japan and the US [3,4]. However, relatively little testing has studied the effects of a composite slab on steel moment connection behavior under cyclic loads. Since an actual steel building Corresponding author. Tel.: +82-2-2290-0312; fax: +82-2-22964145. E-mail address: [email protected] (Y.-J. Kim).
0141-0296/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2004.07.017
has floor slabs that have strength and are bonded to supporting beams from full to partial composite actions, their actual behavior under an earthquake load will be different from that expected from bare steel connections. An illustration of composite beam to column connections subjected to seismic load is shown in Fig. 1. Considering the effect of the concrete slabs, the neutral axis moves toward the top flange when subjected to positive beam bending (concrete slab under compression), and, consequently, this causes the strain on the bottom flange to be much larger than that of the top flange. This strain concentration on the bottom flange may lead to premature failure of the connection, thus reducing the deformation capacity of composite beam connections. Most beam to column connections were designed on the assumption that composite beam connections with slab have the same deformation capacity as bare steel beam connections without slab. The slab effects on the seismic behavior of the connection to date were not yet fully uncovered for the case of a laterally loaded structure. In addition, less testing has been directed towards retrofit methods for
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Nomenclature f’c Ki Mmax Mpc Mps V 1, V 2 W ddt h hsmax hpanel hsp Rlhl
Compressive stress of concrete slab Initial stiffness Maximum flexural moment Plastic moment of composite beam Plastic moment of bare steel beam Vertical displacement at connection by panel zone deformation Absorbed energy Total displacement at loading point Target rotation angle Maximum skeleton rotation angle Panel zone deformation angle Plastic rotation angle corresponding to plastic moment of steel beam Cumulative rotation angle
existing moment connections. This study was also undertaken to investigate the effects of relatively inexpensive and nonintrusive retrofit procedures on connection performance. In this study, which focuses on the presence of a composite slab which may affect the fracture of the beam bottom flange in the beam to column connections which occurred in the Kobe earthquake, a fracture in the composite beam was reproduced experimentally and the effects of the presence of floor slabs were investigated. This study also investigated the effects of the seismic retrofit methods for the enhancement of deformation capacity of existing composite connections designed prior to the Kobe earthquake.
2. Previous studies of composite beam connections Several studies on positive effects of a composite slab were reported. To enhance the ductility of beam to column connections of existing buildings, Chen et al. proposed a simple method which trims the bottom flange of the steel beam slightly [5]. Test results indicated that
Fig. 1. Strain distributions of composite beams.
the concrete slab may contribute another 16% to the connection strength and the plastic rotation angles of composite specimens were more than 0.04 rad. Uang et al. investigated the effectiveness of using RBS and welded haunch for seismic retrofit of pre-Northridge steel moment connections through cyclic testing [6]. A test result showed that the composite slab only increased the beam positive flexural strength and no brittle fracture occurred. In addition, it was found that lateral-torsional buckling was prevented due to the bracing effect of the slab. Civjan et al. stated that the composite beam connections with floor slabs significantly reduced the severity of local and lateral buckling of the beam and corresponded to the higher-achieved overall plastic rotations and higher-peak attained moments compared to bare steel beam connections [7,8]. Other studies of composite connections were reported [9–12]. The tests clearly showed that without early fracture or detrimental effects, the presence of the slab was beneficial to specimen performance by enhancing beam stability and delaying strength degradation. On the contrary, some negative effects were noted with composite specimens. A bare steel and two partially composite specimens designed as 35 and 55% composite were tested by Leon et al. [13] and Hajjar et al. [14]. According to these tests, the presence of a composite slab may increase stress and strain demands on the bottom flange weld and be conjectured that the slab could cause an earlier fracture at the bottom flange weld, thereby degrading connection performance and nullifying any positive effects of the RBS. Okada et al. [15] and Oh et al. [16] conducted experimental and analytical studies on the deformation capacity of the beam to box column connections with and without slabs under cyclic loading. The results of the tests and analysis indicated that the deformation capacity of composite beams was nearly half of that of steel beams
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Fig. 2. Typical Japanese connection.
without slabs, and these results demonstrated earlier fractures of bottom flanges of composite beams as shown in the Kobe Earthquake.
3. Test program 3.1. Test specimen Modern steel moment resisting frames in Japan are usually constructed using square tube columns and wide flange beams known as a ‘‘through-diaphragm connection’’, due to their excellent cross-sectional properties to resist biaxial bending loads as shown in Fig. 2 [2,17]. Top and bottom square tube columns were welded to each other using a 25 mm thick complete penetration diaphragm. The beam flanges were welded to the diaphragms instead of a column flange, and a standard web access hole was used. CJP single bevel groove welds were used to connect the beam flanges to the diaphragm plates, and fillet-welds were used to connect the beam web to the column. The root of the CJP groove welds was located on the interior side of both the top and the bottom flange. Gas metal arc welding (GMAW) with CO2 shielding was used to fabricate the welded joints of test specimens. A solid electrode designated as ER70S-3 was used in all welding. All subassemblages were designed as an exterior connection. Test specimens were constructed using a builtup square tube column Box-45045022 and wide flange beams H-6122021323 with a throughdiaphragm connection (Fig. 3). A relatively strong column was used to ensure that the beams could initiate the development of a plastic hinge mechanism during cyclic loading before damage develops in the column. Also, incorporating a rigid panel zone Box4504502232 may significantly reduce the contribution of panel zone shear deformation to mode of failure. To simulate the actual steel building, four side beams were designed at each column web and each side of loading part, and were fully bonded to the slab using 22 mm diameter shear studs. To make fully composite
Fig. 3.
Typical Specimen (SP-2; Standard).
action between the beam and the slab, two rows of studs were welded to the top surface of the top flange at 200 mm spacing. The concrete slab was designed with common concrete having a design strength of 24 MPa, which was 200 mm thick and 2500 mm width. Also, the concrete slab was provided with longitudinal and transverse steel reinforcement. For reinforcement, D13 (13 mm diameter steel reinforcing bars) having a design strength of 400 MPa were placed transversely at 200 mm spaced over the entire length of the concrete slab and were located at both 30 mm from the top and from the bottom of the slab. In addition, a mould plate for slip prevention between the beam flange and the concrete slab was installed as shown in Fig. 3. The mould plate was chosen so that the slab would contribute more strength to the connection, thereby simulating a more severe strain condition. All steel material is SM490 and a list of material properties used for specimens is given in Table 1. A total of five specimens were manufactured as shown in Table 2, and the specimens are largely divided into two series. 3.1.1. Series I—Conventional specimens Series I consists of a bare steel specimen (SP-1) and a composite specimen (SP-2) with concrete floor slab. They have a conventional type of the weld access hole in their beam-to-column connections (Fig. 2). SP-2 is the standard specimen in this study. The detailed objectives of the tests using series I are: (1) to demonstrate
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Table 1 Steel material characteristics Member
Yield strength (Mpa) Tensile strength (Mpa)
Yield ratio (%)
Elongation (%)
Beam flange (H-6122021323) Beam web (H-6122021323) Column (Box-45045022) Diaphragm (PL-50050025) Panel zone (Box-4504502232)
364 390 395 378 382
68 72 70 67 69
27 26 26 27 27
fractures of the conventional type connections that occurred in the Kobe earthquake; (2) to investigate the deformation capacity of composite beams for fracture; and, (3) to make clear the effects of slabs on the deformation capacity of composite beams, compared with that of bare steel beam connections. 3.1.2. Series II—Retrofit specimens Series II consists of three composite specimens which used the design concept for seismic retrofit methods to improve their deformation capacity. As shown in Fig. 4, the three retrofit schemes include no weld access hole (SP-3), horizontal stiffener (SP-4), and cover plate (SP-5).
532 541 563 565 548
3.1.2.1. No weld access hole scheme (SP-3). Many studies reported the effectiveness of no weld access hole (Fig. 4(a)), which could exhibit excellent seismic performance in steel beam connections with or without concrete floor slabs [3,15,18]. But no testing studied the effects of this retrofit scheme on existing steel moment connection. In general, the existence of a weld access hole is required for CJP welding of the beam flange. However, it was discovered that it was a primary factor for premature brittle fracture due to the influence of the section loss of the web, where the geometry changes abruptly and the joint efficiency (that is, moment transfer efficiency) deceases [19]. To prevent these defects, the no weld access hole retrofit was developed.
Table 2 List of specimens Specimen
Composite or bare steel
Parameter
Connection type
Location of retrofit
SP-1 SP-2 SP-3 SP-4 SP-5
Bare steel Composite Composite Composite Composite
Without slab Standard No weld access hole Horizontal stiffener Bottom plate
Conventional type Conventional type Retrofit type Retrofit type Retrofit type
– – Bottom flange Bottom flange Bottom flange
Fig. 4. Schematic illustrations of retrofit specimens.
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As shown in Fig. 4(b), the welding of a new web plate without an access hole was involved after the removal of the beam web part including a weld access hole and the backing bar. CJP single bevel and square groove welds were used to connect the beam to the new web plate. 3.1.2.2. Horizontal stiffener scheme (SP-4). This scheme was to strengthen the steel beam near the welded connection by welding a horizontal stiffener to the beam bottom flange (Fig. 4(c)). Sugimoto et al. proved that the horizontal stiffener scheme (that is, wing plate connection, see ref. [20]) can successfully prevent premature rupture by neutralizing the negative effect of the weld access hole, backing bars and run-off tabs and exhibit large deformation capacity. Ideally, the strengthened section of the beam would remain primarily elastic during plastic hinge movement of the beam, thereby limiting the stress in the welds. 3.1.2.3. Cover plate scheme (SP-5). The cover plate scheme effectively strengthens the connections so that flexural yielding of the beam at the end of the cover plate can provide the connection ductility [21]. The connection was usually used with either top and bottom cover plate or cover plate at the bottom flange only. In this study, a bottom flange only cover plate was considered primarily as a retrofit to existing moment connections. In addition, the specimen, SP-5, involved slot welding to the beam flange around the weld access hole as shown in Fig. 4(d). The concept of this scheme was to transfer the stress to the bottom cover plate through slot welding. Moreover, it was reinforced with a vertical rib halfway up the beam depth. There was a fundamental difference of retrofit scheme between this connection and the typical cover plate, due to the existence of a slot welding. The concrete slab in existing buildings presents problem for economic considerations and work convenience in seismic retrofit schemes. Unless the concrete slab around the column is removed, it is difficult to retrofit the top flange and its welded joint. Also, because the majority of reported damage occurred in the bottom flange, it was speculated that retrofit only at the bottom flange may be sufficient to significantly improve the performance of the connections. Therefore, all specimens included in Series II were retrofitted at only the bottom flange of the beam.
Fig. 5.
Test setup.
loaded at an assumed pin at its midspan. The test setup is shown in Fig. 5. The loading protocol prescribed a quasi-static cyclic pattern defined in terms of a rotation angle (h) (Fig. 6). In Fig. 6, the target rotation angle h is the rotation angle of only the pure beam with a displacement d that is divided with the distance (L ¼ 3275 mm) from column face to loading point at the total displacement ddt, excluding displacement V1 and displacement V2 by the connection panel zone deformation angle hpanel. The plastic rotation angle hsp corresponding to the plastic moment of bare steel beam, Mps , was calculated, and this rotation angle was incremented at 2hsp , 4hsp , 6hsp , and so on after loading at the elastic rotation. This history was applied to each specimen until failure occurred. The specimen was equipped with displacement transducers and strain gauges to measure deformation contributions of different parts and measure strain distributions, respectively. The measurement position of the strain gauge was shown in Fig. 7. Four sections, A, B, C and D that are respectively 75, 325, 575 and 825 mm distance from the column face are the positions for strain measurement. In particular, section A was located in line of the toe of weld access hole so that the stress situation near the weld access hole could be investigated in detail.
3.2. Test setup The test setup was designed to simulate the boundary conditions of a beam-to-column connection subassembly in a moment resisting frame under typical lateral loading. Thus, the column was assumed to be pin-supported at midstorey points, and the beam was
Fig. 6.
Target rotation angle h.
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Fig. 7.
Position of strain gauges.
4. Test result 4.1. Global behavior and failure mode For the purpose of discussion of the test results, moment versus rotation angle relationships (h) were plotted in Fig. 8, where Mcp (=2173 kN.m) is the calculated plastic moment for composite beams and Mps (=1423 kN.m) is that for bare steel beam without slabs [22,23]. Here, h is the rotation angle of only the pure beam. m, ! and # shows the point where fracture, local buckling and crack initiation of connections occurred, respectively. A summary of experimental results for the five tests is given in Table 3. Photographs of all specimens after testing are shown in Fig. 9. In all specimens with composite beam connections, excluding the bare steel beam connection, the bottom flange failed. The bare steel beam connection (SP-1) showed stable hysteresis behavior and good deformation capacity, however, the connection failed by fracture of the top flange upon the negative bending of
Fig. 8.
6hsp . The initial crack developed at the center of the flange from the toe of the weld access hole. The crack progressed outward along the flange width during successive loading, followed by a sudden break of the entire beam flange. The initial crack propagation was ductile, and the final was brittle (Fig. 9(a)). In standard composite specimen (SP-2), although maximum bending moment increased compared with the bare steel beam specimen (SP-1), deformation capacity appeared small due to the premature brittle failure of the beam bottom flange. The initial ductile crack was located in the area between the weld fusion zone on the outside flange surface and the end of the access hole cut on the inside flange surface, which was similar to the damage which occurred during the Kobe earthquake. After this ductile crack reached its critical size, the brittle fracture occurred as shown in Fig. 9(b). SP-3 specimen failed in the bottom flange beneath the weld of the end of the new web plate abruptly, as shown in Fig. 9(c). The reason might be the unsatisfactory workmanship of the connection. The interruption of the web stiffener near the connections led to the
Moment- rotation angle relationships (m: fracture; !: local buckling; #: crack initiation).
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Table 3 Test results Specimen
Bending direction
Mmax [kN.m]
Ki [kN.m/rad.]
h smax [rad.]
W [KN.m. rad.]
Fracture position
SP-1
Positive (+) Negative () Positive (+) Negative () Positive (+) Negative () Positive (+) Negative () Positive (+) Negative ()
1800 1560 2490 1900 2600 1900 3260 2190 3010 2110
148000 152000 432000 297000 552000 294000 447000 246500 497000 311000
0.052 0.036 0.026 0.040 0.025 0.041 0.037 0.057 0.026 0.038
202
Top flange
134
Bottom flange
172
Bottom flange
405
Bottom flange
195
Bottom flange
SP-2 SP-3 SP-4 SP-5
complete penetration weld between the beam bottom flange and the new web plate becoming difficult and less reliable. Although it was retrofitted with a no weld access hole, the hysteresis characteristic of the specimen SP-3 exhibited nearly similar to the specimen SP-2. Specimen SP-4 exhibited an excellent hysteretic behavior over all other composite specimens through the prevention of stress concentration of the fuse zone including the toe of the weld access hole and heat-affected zone, with the horizontal stiffener remaining elastic. Moreover, upon the negative bending of 6hsp amplitude, the local buckling occurred in the horizontal stiffener end, followed by a degradation of the strength. This finding implies that, in the horizontal stiffener scheme, a strengthened area on the beam near the column face can provide reliable energy dissipation through developing the plastic hinge in the horizontal stiffener end even if the concrete floor slab is added to the connection (Fig. 9(d)). SP-5 induced stress distribution of the bottom flange around the connection by the bottom cover plate. But, eventually, a brittle fracture occurred at the bottom flange outer edge upon the positive bending of 6hsp amplitude (Fig. 9(d)). After the fracture of the connection, the vertical rib was removed for detailed inspection of the failure mode of the connection and it was examined. The crack initiated at the welding part between the diaphragm and beam flange and progressed inward along the flange width through the slot welding. The weld was also examined, and a lack of fusion was also found at the slot welding. Therefore, good workmanship and sound welding quality are still required in order to provide a reliable welded connection. 4.2. Ultimate strength and stiffness
Fig. 9.
Photograph of specimens after testing.
The data in Table 3 showed that the average maximum strength of composite specimens under positive bending was 1.39 times the maximum strength under negative bending. When the concrete slab is under positive bending, the maximum strength of the
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Fig. 10. Definition of maximum skeleton rotation angle and skeleton curve.
composite specimens was 1.58 times the maximum strength of the bare steel beam on average. The ratio of attained positive moments to the calculated plastic capacity of the unreinforced composite specimens (SP-2 and SP-3) was about 1.17 on average. The plastic capacity of the sections of these composite specimens was computed assuming an effective concrete stress of 1.3 f’c [22,23]. The results showed that the computed plastic capacity of the composite section effectively estimated the actual plastic capacity. However, the average maximum positive moment of the reinforced composite specimens (SP-4 and SP-5) was about 1.44 times greater than the estimated plastic capacity. This showed that the contribution of the floor slab added considerable extra strength to the ultimate strength of the beam. In the previous test, the presence of the concrete slab increased the positive flexural strength of the beam by about 10% on average [6]. Composite specimens exhibited a slight increase in initial elastic stiffness and strength to the effect of 10–30% over similar bare steel specimens [7,8]. Another study reported that the peak load resisted by the composite slab was developed between 4 and 17% larger than the corresponding bare steel counterparts [9]. The results of the previous research indicated that the slab did not, in fact, produce a significant increase in the ultimate strength of the beam. Note, however, that the beams used in their studies were relatively deep (i.e. W30 to W36). Thus, the influence of the slab on the strength and stiffness of the girder was assumed to be relatively small when compared to more typical composite beams (i.e. W18 to W24) similar to specimens used in this study. Assuming the thickness of the slab does not decrease linearly with the depth of the beam, it is speculated that shallow beams are potentially more susceptible to composite slab effects because of a great neutral axis shift. The data in Table 3 showed that initial stiffness of composite specimens under positive bending exhibited a significant increase to the effect of 45 to 88% over initial stiffness under positive bending. When the concrete slab was under positive bending, the initial stiffness of the
composite specimens was 3 times that of the bare steel beam on average due to the composite action of the floor slab. The increase of the initial stiffness in composite connections almost developed deformation of the connection in the plastic range rather than the elastic range. For this reason, composite beam connections exhibited a premature fracture due to the early plastic strain concentration. 4.3. Deformation capacity Concepts of a skeleton curve and the Bauschinger curve have been commonly adopted when characterizing the deformation capacity of steel members subjected to load reversal in Japan [24]. As shown in Fig. 10, the moment versus rotation angle curve of steel members subjected to cyclic loading can be decomposed into the skeleton part, the Bauschinger part, and the elastically unloading part. The skeleton curve is constructed from moment versus rotation angle hysteresis loops (i.e. the solid lines in Fig. 10). The fundamental hypothesis behind the obtained skeleton curve for a cyclic loading test is the same as the curve obtained from a monotonic loading test, and the skeleton is thus useful to compare the ductility capacity of hysteresis curves of specimens of various types [24]. The deformation capacity of each specimen was compared using the maximum skeleton rotation angle (hsmax ) which is in positive and negative bending state in moment versus rotation relationships. The skeleton curves under positive bending are shown in Fig. 10, and hsmax is shown in Table 3. hsmax is a value evaluated at 80% (0.8 Mmax) of the ultimate strength of the specimen after it reached ultimate strength in skeleton curves. hsmax of the composite specimen SP-2 under positive bending was 50% when compared to the bare steel specimen (SP-1). Due to the composite action of the slab, SP-2 had smaller ductility than that of bare steel specimen, SP-1. hsmax of Series II specimens SP-3, SP-4, and SP-5 with retrofit details were 48, 71 and 50% when compared to specimen SP-1. hsmax of SP-4 was 42% greater than that of the typical composite
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Fig. 11. Plastic neutral axis and strain distributions.
specimen SP-2. When the slab was under tension, the deformation capacity of the horizontal stiffener specimen SP-4 was also higher than that of the other composite specimens, and even that of the bare steel beam specimen SP-1 without a slab. Absorbed energy was also evaluated. Absorbed energy (W) of all specimens is shown in Table 3. the energy of the composite specimens was 66–201% when compared to specimen SP-1 without a slab. Especially, specimen SP-4 was two times higher than SP-1.
Test results exhibited that, while composite connections developed higher moment than bare steel connections, composite connections needed more local deformation capacity than bare steel connections need. The effect of the floor slabs was investigated in this section based on the results of the measure strain distributions.
It was found that the strain on the bottom flange was several times larger than that of the top flange in the previous studies [5,15]. Focusing on these results, the plastic neutral axis of bending in the beams was estimated from strain behavior obtained at section A. The plastic neutral axis at section A of all specimens under cyclic load reversals is shown in Fig. 11(a). The position of the neutral axis can be obtained by using the distance ratio of the height of the beam depth and the strain increment of both bottom and top beam flange. As shown in Fig. 11(a), the plastic neutral axis of SP-1 is located in the vicinity of the center of the beam web, but that of SP-2 is shown to be near the top flange level. Consequently, this caused the strain on the bottom flange to be much larger than that of the top flange (Fig. 11(b)). As shown in Fig. 11(b), while the strain of SP-1 showed a symmetric balance for the top and bottom flange, the test result of SP-2 exhibited that it was concentrated in the bottom flange. The strain of the bottom flange for SP-1 and SP-2 under positive bending is shown in Fig. 11(c). At the same cumulative rotation angle, Rlhl ¼ 0:1 rad, the strain of the composite beam (SP-2) was 4.4 times higher than that of the bare steel beam (SP-1).
5.1.1. Strain concentration and plastic neutral axis The concrete floor slabs may increase the positive flexural capacity, raising the beam neutral axis and, therefore, creating a large strain in the bottom flange.
5.1.2. Moment transfer efficiency of beam web A square tube column had two webs at each side but no web at the center where the beam web was connected. Therefore, the web of the box column was signifi-
5. Discussion 5.1. Composite effects of slabs (Series I—Conventional specimens)
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Fig. 12. Deformation of column flange and strain distribution of connection.
cantly less effective in transferring bending moment due to the out-of-deformation of the column flanges and the loss of the web sections by the weld access hole (Fig. 12(a)). Stress increase at the end of the beam is influenced by the efficiency in transmitting the stress in the web of beam through beam to column joint. Moreover, it was reported that the beam near the connection developed a smaller strength than the beam away from the connection due to the decrease of the web moment transfer as shown in Fig. 12(b) [19]. To investigate the effect of web moment transfer, moment versus curvature relationship was used. The curvature was obtained from the gross sectional average strain of a beam flange. Next, the skeleton curve of a curvature constructed from moment versus curvature relationships and moment versus monotonic curvature relationships at each section under positive bending is shown in Fig. 13. In Fig. 13(a), the moment of section A in bare steel specimen SP-1 appeared 8, 11, and 13% slightly smaller than sections B, C and D at the same curvature, respectively. This means that the web connection did not transfer its full share of the beam moment, resulting in strain concentration of the beam flange. The moment of section A in composite specimen SP-2 exhibited 35, 30, and 36% significantly smaller than sections B, C and D at the same curvature, respectively (Fig. 13(b)). The decrement of the strength of SP-2 was about 3 times of that of SP-1, and this exhibited that composite connection decreased the web moment transfer compared to bare steel connection due to the composite effect of the slab. Therefore, the deformation was concentrated near the beam bottom flange where the constraint was minimal. More detailed studies are required for these results.
5.1.3. Reduction of plastic zone Fig. 14(a) shows the distributions of the curvature in the beam’s longitudinal direction at the completion of the peak moment, and it was assumed that there was almost no curvature change from section D to loading point and it accordingly was simplified into a linear line. Fig. 14(b) shows the ratio of the curvature of each section to the summation of those for all sections. SP-1 had a curvature distribution that changed gradually from section A to section D. But in SP-2, the curvature distribution from section A to section B changed exceedingly and the curvatures from section B to section D were shown to be very small compared with section A. These results also can be seen clearly in Fig. 14(b). Fig. 15 shows the absorbed energy in each section of all specimens. The energy obtained from moment versus curvature hysteresis loops. This plot exhibited that, in the case of SP-2, the ratio of the energy absorbed in section A to that in all sections was near 80%, while, in the case of SP-1, it was a ratio of 59%. Also, while the energy of SP-1 showed a symmetric balance between the positive and negative bending, the test result of SP2 exhibited that it was concentrated in the positive bending direction. These results indicated that, for bare steel connections, the deformation was relatively well distributed over the entire section, but, for composite connections, it had higher local ductility demand in a narrow plasitified zone of the beam. In other words, it can be understood that the plastic zone of composite beams was reduced while bare steel beams secured a wide plastic zone. 5.2. Effect of seismic retrofit (Series II—Retrofit specimens) In previous studies, no weld access hole connections showed outstanding performance in laboratory testing due to the high efficiency of web moment transfer [15,19]. While typical details for new construction used no weld access hole scheme in both the top and bottom web, the retrofit of Specimen SP-3 used it only in the bottom web. In Fig. 13(c), the moment of section A in specimen SP-3 exhibited 18, 8, and 7% smaller than sections B, C, and D, at the same curvature, respectively. The decrement of the strength of SP-3 was significantly smaller than that of SP-2, exhibiting the ratio of 30, 36 and 35%. The decrement of the strength of SP-3 was slightly smaller than that of SP-2. To directly compare moment transfer efficiency of SP-3 with SP-2, moment versus curvature relationship and strain value was plotted in Fig. 16. Fig. 16 indicated that, at the same curvature level, SP-2 and SP-3 specimens had different flexural strength at the critical region, section A. For example, at the curvature of 48 (106/mm), the
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Fig. 13. Moment versus curvature relationships.
strength of the beam for SP-1 already reached 2530 KN.m, while, for SP-2, the strength of the beam was only 2154 KN.m. As shown in Fig. 17, in the same cumulative rotation angle, the strain on the bottom flange of SP-3 was much smaller than that of SP-2 at section A, while it was slightly larger than that of SP-1. Also, Fig. 14 showed that in the case of SP-3, the ratio of the curvature in section A to that in all sections was about 59%, while, in the case of SP-2, it was a ratio of 83%. In addition, the plastic neutral axis of SP-3 was
Fig. 14. Distribution of curvature along beam length.
Fig. 15. Absorbed energy in each section.
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Fig. 16. Effect of no weld access hole scheme.
located slightly lower than that of SP-2 at the same cumulative rotation angle (Fig. 11(a)). These results indicated that the no weld access hole scheme was effective in transferring the web moment even in composite connections, so that it developed less ductility demand on the beam bottom flanges. But, the notch caused by the unsatisfactory workmanship developed and, consequently, SP-3 failed in the bottom flange beneath the weld of the new web plate end prior to developing a sufficient deformation capacity, suggesting that the reliability of the workmanship was needed in doing field constructions of retrofit schemes. The horizontal stiffener specimen (SP-4) improved the ductility of composite connections and showed stable hysteretic characteristics compared to the standard composite specimen (SP-2). This specimen exhibited flange local buckling and reached failure due to a ductile crack initiated at the tip of the horizontal stiffener. Therefore, SP-4 also improved the energy absorption capacity of composite beam connections when compared to specimen SP-2. Fig.13(d) shows that for SP-4, section A and section B of the beam near the column face stiffened by the horizontal stiffener remained in elastic range or deformed only slightly even under ultimate state. The force demand in the existing bottom flange groove weld and the toe of the weld access hole was significantly reduced to a reasonable level, although the plastic neutral axis of this specimen moved forward to the top flange but differed from that of specimen SP-2, as shown in Fig. 11(a). Fig. 14(a) shows the distributions of curvature in the beam’s longitudinal direction at completion of the peak moment. This plot showed that for SP-4, the curvature was small near the connections and increased from the horizontal stiffener end. In addition, in Fig. 14(b), the ratio of the curvatures in sections A, B, C, and D to that in all sections was 7, 8, 50 and 35%, respectively. Similar results can also be seen through the absorbed energy of each section. As shown in Fig. 15, the energy of SP-2 concentrated in section A of 80% compared with all sections. But, SP-4 absorbed energy in section C and section D of 50% compared with all sections. These results may illustrate that moment connections retrofitted using the horizontal stiffener moved the
Fig. 17. Strain distribution.
plastic hinge away from the column face and, thus, achieved a more reliable connection performance even in composite connections. Although the cover plate specimen (SP-5) provides a strengthened zone in the vicinity of the connection that results in a lower level of stress/strain on the weld as shown in Figs.13(e), 14 and 17, the proposed connection scheme exhibited poor performance. As mentioned in the tests results, good workmanship and sound welding quality are still required in order to provide a reliable welded connection. 6. Conclusion The presence of the concrete slabs had an influence on the neutral axis location, the web moment transfer, and the plastic zone of the connection, and these influences caused the strain on the bottom flange to be much larger than that of the top flange. Consequently, the deformation capacity of a composite connection is about half that of a bare steel connection. These results indicated that the detrimental slab effects should be considered in the seismic design of the connection. The no weld access hole detail (SP-3) was effective in the transferring of the web moment even in composite connections, such that it developed less ductility demand on the beam bottom flanges. The cover plate specimen (SP-5) provided a strengthened zone in the vicinity of the connection that can be expected to result in a lower level of stress/strain on the weld. Nevertheless, these proposed connections exhibited poor performance, suggesting that reliability of the workmanship and welding quality were required to ensure the ductile performance of the connection. The horizontal stiffener details (SP-4) showed that an enlarged plastic zone can be achieved and the deformation capacity can be improved. This retrofit scheme demonstrated very good potential in improving the ductility of composite beam connections in existing buildings without removing the floor slabs. Moreover, by using the horizontal stiffener in retrofit of the existing building, workmanship as well as excellent performance
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can be expected at the same time. This is because by using the horizontal stiffener retrofit scheme, downward welding is possible and the design scheme is very simple. Acknowledgements The author gratefully acknowledges the support of STRESS and POSCO, for financially supporting this research. References [1] Miller DK. Lessons learned from the Northridge earthquake. Eng Struct 1998;20(4-6):249–60. [2] Nakashima M, Inoue K, Tada M. Classification of damage to steel buildings observed in the 1995 Hyogoken-Nanbu earthquake. Eng Struct 1998;20(4-6):271–81. [3] FEMA. Recommended seismic Design Criteria For New Steel moment Buildings. FEMA-350. Washington DC: Federal Emergency Management Agency, 2000. [4] Nakashima M, Suita K, Morisako K, Maruoka Y. Tests of welded beam-column subassemblies. I: Global behavior. J Struct Eng, ASCE 1998;124(11):1236–44. [5] Chen S-J, Tsao YC, Chao C. Enhancement of ductility of existing seismic steel moment connections. J Struct Eng, ASCE 2001;127(5):538–45. [6] Uang C-M, Kent Yu Q-S, Noel S, Gross J. Cyclic testing of steel moment connections rehabilitated with RBS or welded haunch. J Struct Eng, ASCE 2000;126(1):57–68. [7] Civjan SA, Engelhardt MD, Gross JL. Retrofit of pre-Northridge moment-resisting connections. J Struct Eng, ASCE 2000;126(4):445–52. [8] Civjan SA, Engelhardt MD, Gross JL. Slab effects in SMRF retrofit connection tests. J Struct Eng, ASCE 2001;127(3):230–7. [9] Jones SL, Fry GT, Engelhardt MD. Experimental evaluation of cyclically loaded reduced beam section moment connections. J Struct Eng, ASCE 2001;128(4):441–51. [10] Roeder CW. General issues influencing connection performance. J Struct Eng, ASCE 2001;128(4):420–8.
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[11] Roeder CW. Connection performance for seismic design of steel moment frames. J Struct Eng , ASCE 2001;128(4):517–25. [12] FEMA. Connection performance state of art report. FEMA355D. Washington, DC. Federal Emergency Management Agency, 2000. [13] Leon RT, Hajjar JF, Gustafson MA. Seismic response of composite moment-resisting connections. I: Performance. J Struct Eng, ASCE 1998;124(8):868–76. [14] Hajjar JF, Leon RT, Gustafson MA, Shield CK. Seismic response of composite moment-resisting connections. II: Behavior. J Struct Eng, ASCE 1998;124(8):877–85. [15] Okada K, et al. Experimental study on deformation capacity of the composite beams. In: Proceeding of 6th Pacific Structural Steel Conference (PSSC), Beijing, China, 2001:233–8 [16] Oh SH, Joh CB, Chang IH, Chen W-F. Slab effects on the seismic behavior of steel moment connections. In: Proceedings of 6th Pacific Struct Steel Conference (PSSC), Beijing, China, 2001:539–44 [17] Kurobane Y. Connections in tubular structures. Prog Struct Eng and Mater 2002;4(1):35–45. [18] Suita K, Tamura T, Morita S, Nakashima M, Engelhardt MD. Plastic rotation capacity of steel beam to column connections using a reduced beam section and no weld access hole design— Full scale tests for improved steel beam to column subassemblies—part 1. J of Struct and Const Eng, AIJ 1999;526(12) :177–84 (in Japanese). [19] Okada K, Matsumoto Y, Yamada S. Evaluation of effect of joint efficiency at beam-to-column connection on ductility capacity of steel beams. J of Struct and Const Eng, AIJ 2003;588(6):131–8 (in Japanese). [20] Sugimoto H, Morita K, Takahashi Y. Prediction of failure mode of hunched beam to SHS column connections. Steel Struct, KSSC 2001;1(3):185–99. [21] Kim T, Whittaker AS, Gilani ASJ, Berto VV, Takhirov SM. Cover-Plate and Flange-Plate Steel Moment-Resisting Connections. J Struct Eng, ASCE 2002;128(4):474–82. [22] Recommendation for limit state design of steel structure, Architectural Institute of Japan, 1998. [23] Design recommendations for composite constructions, Architectural Institute of Japan, 1985 [24] Kato B, Akiyama H. Predictable properties of material under incremental cyclic loading. Lisbon: IABSE Preliminary Publication; 1973.
Seismic behavior and retrofit of steel moment connections considering slab effects Young-Ju Kim a,, Sang-Hoon Oh b, Tae-Sup Moon a a b
Department of Architectural engineering, Hanyang University,17, Haengdang-dong, Seongdong-gu, Seoul,133-791, South Korea Research Institute of Industrial Science & Technology, 79, Youngcheon Dongtan, Hwasung, Kyoungkido, 445-810, South Korea Received 29 January 2004; received in revised form 19 July 2004; accepted 27 July 2004
Abstract A series of five full-scale subassemblages were tested to investigate the contribution of the slabs and the effects of three types of retrofit methods, no weld access hole, horizontal stiffener, and cover plate. The test included one bare steel specimen and four composite specimens with floor slab. The results that emphasize the influence of the composite slab on connection behavior and specific comments on the retrofit scheme are presented. The test result indicated that the strains near the bottom flange of the composite beam connections were several times larger than those of the bare steel beam connections, resulting in a higher potential of fracture. Therefore, the slab effects are detrimental to the seismic behavior of the connection and should be considered in the design. Horizontal stiffener detail of three retrofit schemes demonstrated very good potential in improving the ductility of composite connections in existing buildings. # 2004 Elsevier Ltd. All rights reserved. Keywords: Slab effect; Retrofit; Deformation capacity; Strain concentration; Weld access hole; Horizontal stiffener
1. Introduction During the Northridge earthquake in California in 1994 and the Kobe earthquake in 1995, many steel buildings suffered from fracturing in the welded moment connections [1,2]. The performance of steel buildings during these earthquakes has raised questions regarding the reliability of the current connection method. The current connection design may be susceptible to failure in an earthquake. Investigation of these fractured connections indicated that most of the fractures occurred at the bottom flange. After the earthquake, many different types of connection methods have been suggested to solve the problem both in Japan and the US [3,4]. However, relatively little testing has studied the effects of a composite slab on steel moment connection behavior under cyclic loads. Since an actual steel building Corresponding author. Tel.: +82-2-2290-0312; fax: +82-2-22964145. E-mail address: [email protected] (Y.-J. Kim).
0141-0296/$ - see front matter # 2004 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2004.07.017
has floor slabs that have strength and are bonded to supporting beams from full to partial composite actions, their actual behavior under an earthquake load will be different from that expected from bare steel connections. An illustration of composite beam to column connections subjected to seismic load is shown in Fig. 1. Considering the effect of the concrete slabs, the neutral axis moves toward the top flange when subjected to positive beam bending (concrete slab under compression), and, consequently, this causes the strain on the bottom flange to be much larger than that of the top flange. This strain concentration on the bottom flange may lead to premature failure of the connection, thus reducing the deformation capacity of composite beam connections. Most beam to column connections were designed on the assumption that composite beam connections with slab have the same deformation capacity as bare steel beam connections without slab. The slab effects on the seismic behavior of the connection to date were not yet fully uncovered for the case of a laterally loaded structure. In addition, less testing has been directed towards retrofit methods for
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Nomenclature f’c Ki Mmax Mpc Mps V 1, V 2 W ddt h hsmax hpanel hsp Rlhl
Compressive stress of concrete slab Initial stiffness Maximum flexural moment Plastic moment of composite beam Plastic moment of bare steel beam Vertical displacement at connection by panel zone deformation Absorbed energy Total displacement at loading point Target rotation angle Maximum skeleton rotation angle Panel zone deformation angle Plastic rotation angle corresponding to plastic moment of steel beam Cumulative rotation angle
existing moment connections. This study was also undertaken to investigate the effects of relatively inexpensive and nonintrusive retrofit procedures on connection performance. In this study, which focuses on the presence of a composite slab which may affect the fracture of the beam bottom flange in the beam to column connections which occurred in the Kobe earthquake, a fracture in the composite beam was reproduced experimentally and the effects of the presence of floor slabs were investigated. This study also investigated the effects of the seismic retrofit methods for the enhancement of deformation capacity of existing composite connections designed prior to the Kobe earthquake.
2. Previous studies of composite beam connections Several studies on positive effects of a composite slab were reported. To enhance the ductility of beam to column connections of existing buildings, Chen et al. proposed a simple method which trims the bottom flange of the steel beam slightly [5]. Test results indicated that
Fig. 1. Strain distributions of composite beams.
the concrete slab may contribute another 16% to the connection strength and the plastic rotation angles of composite specimens were more than 0.04 rad. Uang et al. investigated the effectiveness of using RBS and welded haunch for seismic retrofit of pre-Northridge steel moment connections through cyclic testing [6]. A test result showed that the composite slab only increased the beam positive flexural strength and no brittle fracture occurred. In addition, it was found that lateral-torsional buckling was prevented due to the bracing effect of the slab. Civjan et al. stated that the composite beam connections with floor slabs significantly reduced the severity of local and lateral buckling of the beam and corresponded to the higher-achieved overall plastic rotations and higher-peak attained moments compared to bare steel beam connections [7,8]. Other studies of composite connections were reported [9–12]. The tests clearly showed that without early fracture or detrimental effects, the presence of the slab was beneficial to specimen performance by enhancing beam stability and delaying strength degradation. On the contrary, some negative effects were noted with composite specimens. A bare steel and two partially composite specimens designed as 35 and 55% composite were tested by Leon et al. [13] and Hajjar et al. [14]. According to these tests, the presence of a composite slab may increase stress and strain demands on the bottom flange weld and be conjectured that the slab could cause an earlier fracture at the bottom flange weld, thereby degrading connection performance and nullifying any positive effects of the RBS. Okada et al. [15] and Oh et al. [16] conducted experimental and analytical studies on the deformation capacity of the beam to box column connections with and without slabs under cyclic loading. The results of the tests and analysis indicated that the deformation capacity of composite beams was nearly half of that of steel beams
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1995
Fig. 2. Typical Japanese connection.
without slabs, and these results demonstrated earlier fractures of bottom flanges of composite beams as shown in the Kobe Earthquake.
3. Test program 3.1. Test specimen Modern steel moment resisting frames in Japan are usually constructed using square tube columns and wide flange beams known as a ‘‘through-diaphragm connection’’, due to their excellent cross-sectional properties to resist biaxial bending loads as shown in Fig. 2 [2,17]. Top and bottom square tube columns were welded to each other using a 25 mm thick complete penetration diaphragm. The beam flanges were welded to the diaphragms instead of a column flange, and a standard web access hole was used. CJP single bevel groove welds were used to connect the beam flanges to the diaphragm plates, and fillet-welds were used to connect the beam web to the column. The root of the CJP groove welds was located on the interior side of both the top and the bottom flange. Gas metal arc welding (GMAW) with CO2 shielding was used to fabricate the welded joints of test specimens. A solid electrode designated as ER70S-3 was used in all welding. All subassemblages were designed as an exterior connection. Test specimens were constructed using a builtup square tube column Box-45045022 and wide flange beams H-6122021323 with a throughdiaphragm connection (Fig. 3). A relatively strong column was used to ensure that the beams could initiate the development of a plastic hinge mechanism during cyclic loading before damage develops in the column. Also, incorporating a rigid panel zone Box4504502232 may significantly reduce the contribution of panel zone shear deformation to mode of failure. To simulate the actual steel building, four side beams were designed at each column web and each side of loading part, and were fully bonded to the slab using 22 mm diameter shear studs. To make fully composite
Fig. 3.
Typical Specimen (SP-2; Standard).
action between the beam and the slab, two rows of studs were welded to the top surface of the top flange at 200 mm spacing. The concrete slab was designed with common concrete having a design strength of 24 MPa, which was 200 mm thick and 2500 mm width. Also, the concrete slab was provided with longitudinal and transverse steel reinforcement. For reinforcement, D13 (13 mm diameter steel reinforcing bars) having a design strength of 400 MPa were placed transversely at 200 mm spaced over the entire length of the concrete slab and were located at both 30 mm from the top and from the bottom of the slab. In addition, a mould plate for slip prevention between the beam flange and the concrete slab was installed as shown in Fig. 3. The mould plate was chosen so that the slab would contribute more strength to the connection, thereby simulating a more severe strain condition. All steel material is SM490 and a list of material properties used for specimens is given in Table 1. A total of five specimens were manufactured as shown in Table 2, and the specimens are largely divided into two series. 3.1.1. Series I—Conventional specimens Series I consists of a bare steel specimen (SP-1) and a composite specimen (SP-2) with concrete floor slab. They have a conventional type of the weld access hole in their beam-to-column connections (Fig. 2). SP-2 is the standard specimen in this study. The detailed objectives of the tests using series I are: (1) to demonstrate
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Table 1 Steel material characteristics Member
Yield strength (Mpa) Tensile strength (Mpa)
Yield ratio (%)
Elongation (%)
Beam flange (H-6122021323) Beam web (H-6122021323) Column (Box-45045022) Diaphragm (PL-50050025) Panel zone (Box-4504502232)
364 390 395 378 382
68 72 70 67 69
27 26 26 27 27
fractures of the conventional type connections that occurred in the Kobe earthquake; (2) to investigate the deformation capacity of composite beams for fracture; and, (3) to make clear the effects of slabs on the deformation capacity of composite beams, compared with that of bare steel beam connections. 3.1.2. Series II—Retrofit specimens Series II consists of three composite specimens which used the design concept for seismic retrofit methods to improve their deformation capacity. As shown in Fig. 4, the three retrofit schemes include no weld access hole (SP-3), horizontal stiffener (SP-4), and cover plate (SP-5).
532 541 563 565 548
3.1.2.1. No weld access hole scheme (SP-3). Many studies reported the effectiveness of no weld access hole (Fig. 4(a)), which could exhibit excellent seismic performance in steel beam connections with or without concrete floor slabs [3,15,18]. But no testing studied the effects of this retrofit scheme on existing steel moment connection. In general, the existence of a weld access hole is required for CJP welding of the beam flange. However, it was discovered that it was a primary factor for premature brittle fracture due to the influence of the section loss of the web, where the geometry changes abruptly and the joint efficiency (that is, moment transfer efficiency) deceases [19]. To prevent these defects, the no weld access hole retrofit was developed.
Table 2 List of specimens Specimen
Composite or bare steel
Parameter
Connection type
Location of retrofit
SP-1 SP-2 SP-3 SP-4 SP-5
Bare steel Composite Composite Composite Composite
Without slab Standard No weld access hole Horizontal stiffener Bottom plate
Conventional type Conventional type Retrofit type Retrofit type Retrofit type
– – Bottom flange Bottom flange Bottom flange
Fig. 4. Schematic illustrations of retrofit specimens.
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As shown in Fig. 4(b), the welding of a new web plate without an access hole was involved after the removal of the beam web part including a weld access hole and the backing bar. CJP single bevel and square groove welds were used to connect the beam to the new web plate. 3.1.2.2. Horizontal stiffener scheme (SP-4). This scheme was to strengthen the steel beam near the welded connection by welding a horizontal stiffener to the beam bottom flange (Fig. 4(c)). Sugimoto et al. proved that the horizontal stiffener scheme (that is, wing plate connection, see ref. [20]) can successfully prevent premature rupture by neutralizing the negative effect of the weld access hole, backing bars and run-off tabs and exhibit large deformation capacity. Ideally, the strengthened section of the beam would remain primarily elastic during plastic hinge movement of the beam, thereby limiting the stress in the welds. 3.1.2.3. Cover plate scheme (SP-5). The cover plate scheme effectively strengthens the connections so that flexural yielding of the beam at the end of the cover plate can provide the connection ductility [21]. The connection was usually used with either top and bottom cover plate or cover plate at the bottom flange only. In this study, a bottom flange only cover plate was considered primarily as a retrofit to existing moment connections. In addition, the specimen, SP-5, involved slot welding to the beam flange around the weld access hole as shown in Fig. 4(d). The concept of this scheme was to transfer the stress to the bottom cover plate through slot welding. Moreover, it was reinforced with a vertical rib halfway up the beam depth. There was a fundamental difference of retrofit scheme between this connection and the typical cover plate, due to the existence of a slot welding. The concrete slab in existing buildings presents problem for economic considerations and work convenience in seismic retrofit schemes. Unless the concrete slab around the column is removed, it is difficult to retrofit the top flange and its welded joint. Also, because the majority of reported damage occurred in the bottom flange, it was speculated that retrofit only at the bottom flange may be sufficient to significantly improve the performance of the connections. Therefore, all specimens included in Series II were retrofitted at only the bottom flange of the beam.
Fig. 5.
Test setup.
loaded at an assumed pin at its midspan. The test setup is shown in Fig. 5. The loading protocol prescribed a quasi-static cyclic pattern defined in terms of a rotation angle (h) (Fig. 6). In Fig. 6, the target rotation angle h is the rotation angle of only the pure beam with a displacement d that is divided with the distance (L ¼ 3275 mm) from column face to loading point at the total displacement ddt, excluding displacement V1 and displacement V2 by the connection panel zone deformation angle hpanel. The plastic rotation angle hsp corresponding to the plastic moment of bare steel beam, Mps , was calculated, and this rotation angle was incremented at 2hsp , 4hsp , 6hsp , and so on after loading at the elastic rotation. This history was applied to each specimen until failure occurred. The specimen was equipped with displacement transducers and strain gauges to measure deformation contributions of different parts and measure strain distributions, respectively. The measurement position of the strain gauge was shown in Fig. 7. Four sections, A, B, C and D that are respectively 75, 325, 575 and 825 mm distance from the column face are the positions for strain measurement. In particular, section A was located in line of the toe of weld access hole so that the stress situation near the weld access hole could be investigated in detail.
3.2. Test setup The test setup was designed to simulate the boundary conditions of a beam-to-column connection subassembly in a moment resisting frame under typical lateral loading. Thus, the column was assumed to be pin-supported at midstorey points, and the beam was
Fig. 6.
Target rotation angle h.
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Fig. 7.
Position of strain gauges.
4. Test result 4.1. Global behavior and failure mode For the purpose of discussion of the test results, moment versus rotation angle relationships (h) were plotted in Fig. 8, where Mcp (=2173 kN.m) is the calculated plastic moment for composite beams and Mps (=1423 kN.m) is that for bare steel beam without slabs [22,23]. Here, h is the rotation angle of only the pure beam. m, ! and # shows the point where fracture, local buckling and crack initiation of connections occurred, respectively. A summary of experimental results for the five tests is given in Table 3. Photographs of all specimens after testing are shown in Fig. 9. In all specimens with composite beam connections, excluding the bare steel beam connection, the bottom flange failed. The bare steel beam connection (SP-1) showed stable hysteresis behavior and good deformation capacity, however, the connection failed by fracture of the top flange upon the negative bending of
Fig. 8.
6hsp . The initial crack developed at the center of the flange from the toe of the weld access hole. The crack progressed outward along the flange width during successive loading, followed by a sudden break of the entire beam flange. The initial crack propagation was ductile, and the final was brittle (Fig. 9(a)). In standard composite specimen (SP-2), although maximum bending moment increased compared with the bare steel beam specimen (SP-1), deformation capacity appeared small due to the premature brittle failure of the beam bottom flange. The initial ductile crack was located in the area between the weld fusion zone on the outside flange surface and the end of the access hole cut on the inside flange surface, which was similar to the damage which occurred during the Kobe earthquake. After this ductile crack reached its critical size, the brittle fracture occurred as shown in Fig. 9(b). SP-3 specimen failed in the bottom flange beneath the weld of the end of the new web plate abruptly, as shown in Fig. 9(c). The reason might be the unsatisfactory workmanship of the connection. The interruption of the web stiffener near the connections led to the
Moment- rotation angle relationships (m: fracture; !: local buckling; #: crack initiation).
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Table 3 Test results Specimen
Bending direction
Mmax [kN.m]
Ki [kN.m/rad.]
h smax [rad.]
W [KN.m. rad.]
Fracture position
SP-1
Positive (+) Negative () Positive (+) Negative () Positive (+) Negative () Positive (+) Negative () Positive (+) Negative ()
1800 1560 2490 1900 2600 1900 3260 2190 3010 2110
148000 152000 432000 297000 552000 294000 447000 246500 497000 311000
0.052 0.036 0.026 0.040 0.025 0.041 0.037 0.057 0.026 0.038
202
Top flange
134
Bottom flange
172
Bottom flange
405
Bottom flange
195
Bottom flange
SP-2 SP-3 SP-4 SP-5
complete penetration weld between the beam bottom flange and the new web plate becoming difficult and less reliable. Although it was retrofitted with a no weld access hole, the hysteresis characteristic of the specimen SP-3 exhibited nearly similar to the specimen SP-2. Specimen SP-4 exhibited an excellent hysteretic behavior over all other composite specimens through the prevention of stress concentration of the fuse zone including the toe of the weld access hole and heat-affected zone, with the horizontal stiffener remaining elastic. Moreover, upon the negative bending of 6hsp amplitude, the local buckling occurred in the horizontal stiffener end, followed by a degradation of the strength. This finding implies that, in the horizontal stiffener scheme, a strengthened area on the beam near the column face can provide reliable energy dissipation through developing the plastic hinge in the horizontal stiffener end even if the concrete floor slab is added to the connection (Fig. 9(d)). SP-5 induced stress distribution of the bottom flange around the connection by the bottom cover plate. But, eventually, a brittle fracture occurred at the bottom flange outer edge upon the positive bending of 6hsp amplitude (Fig. 9(d)). After the fracture of the connection, the vertical rib was removed for detailed inspection of the failure mode of the connection and it was examined. The crack initiated at the welding part between the diaphragm and beam flange and progressed inward along the flange width through the slot welding. The weld was also examined, and a lack of fusion was also found at the slot welding. Therefore, good workmanship and sound welding quality are still required in order to provide a reliable welded connection. 4.2. Ultimate strength and stiffness
Fig. 9.
Photograph of specimens after testing.
The data in Table 3 showed that the average maximum strength of composite specimens under positive bending was 1.39 times the maximum strength under negative bending. When the concrete slab is under positive bending, the maximum strength of the
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Fig. 10. Definition of maximum skeleton rotation angle and skeleton curve.
composite specimens was 1.58 times the maximum strength of the bare steel beam on average. The ratio of attained positive moments to the calculated plastic capacity of the unreinforced composite specimens (SP-2 and SP-3) was about 1.17 on average. The plastic capacity of the sections of these composite specimens was computed assuming an effective concrete stress of 1.3 f’c [22,23]. The results showed that the computed plastic capacity of the composite section effectively estimated the actual plastic capacity. However, the average maximum positive moment of the reinforced composite specimens (SP-4 and SP-5) was about 1.44 times greater than the estimated plastic capacity. This showed that the contribution of the floor slab added considerable extra strength to the ultimate strength of the beam. In the previous test, the presence of the concrete slab increased the positive flexural strength of the beam by about 10% on average [6]. Composite specimens exhibited a slight increase in initial elastic stiffness and strength to the effect of 10–30% over similar bare steel specimens [7,8]. Another study reported that the peak load resisted by the composite slab was developed between 4 and 17% larger than the corresponding bare steel counterparts [9]. The results of the previous research indicated that the slab did not, in fact, produce a significant increase in the ultimate strength of the beam. Note, however, that the beams used in their studies were relatively deep (i.e. W30 to W36). Thus, the influence of the slab on the strength and stiffness of the girder was assumed to be relatively small when compared to more typical composite beams (i.e. W18 to W24) similar to specimens used in this study. Assuming the thickness of the slab does not decrease linearly with the depth of the beam, it is speculated that shallow beams are potentially more susceptible to composite slab effects because of a great neutral axis shift. The data in Table 3 showed that initial stiffness of composite specimens under positive bending exhibited a significant increase to the effect of 45 to 88% over initial stiffness under positive bending. When the concrete slab was under positive bending, the initial stiffness of the
composite specimens was 3 times that of the bare steel beam on average due to the composite action of the floor slab. The increase of the initial stiffness in composite connections almost developed deformation of the connection in the plastic range rather than the elastic range. For this reason, composite beam connections exhibited a premature fracture due to the early plastic strain concentration. 4.3. Deformation capacity Concepts of a skeleton curve and the Bauschinger curve have been commonly adopted when characterizing the deformation capacity of steel members subjected to load reversal in Japan [24]. As shown in Fig. 10, the moment versus rotation angle curve of steel members subjected to cyclic loading can be decomposed into the skeleton part, the Bauschinger part, and the elastically unloading part. The skeleton curve is constructed from moment versus rotation angle hysteresis loops (i.e. the solid lines in Fig. 10). The fundamental hypothesis behind the obtained skeleton curve for a cyclic loading test is the same as the curve obtained from a monotonic loading test, and the skeleton is thus useful to compare the ductility capacity of hysteresis curves of specimens of various types [24]. The deformation capacity of each specimen was compared using the maximum skeleton rotation angle (hsmax ) which is in positive and negative bending state in moment versus rotation relationships. The skeleton curves under positive bending are shown in Fig. 10, and hsmax is shown in Table 3. hsmax is a value evaluated at 80% (0.8 Mmax) of the ultimate strength of the specimen after it reached ultimate strength in skeleton curves. hsmax of the composite specimen SP-2 under positive bending was 50% when compared to the bare steel specimen (SP-1). Due to the composite action of the slab, SP-2 had smaller ductility than that of bare steel specimen, SP-1. hsmax of Series II specimens SP-3, SP-4, and SP-5 with retrofit details were 48, 71 and 50% when compared to specimen SP-1. hsmax of SP-4 was 42% greater than that of the typical composite
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Fig. 11. Plastic neutral axis and strain distributions.
specimen SP-2. When the slab was under tension, the deformation capacity of the horizontal stiffener specimen SP-4 was also higher than that of the other composite specimens, and even that of the bare steel beam specimen SP-1 without a slab. Absorbed energy was also evaluated. Absorbed energy (W) of all specimens is shown in Table 3. the energy of the composite specimens was 66–201% when compared to specimen SP-1 without a slab. Especially, specimen SP-4 was two times higher than SP-1.
Test results exhibited that, while composite connections developed higher moment than bare steel connections, composite connections needed more local deformation capacity than bare steel connections need. The effect of the floor slabs was investigated in this section based on the results of the measure strain distributions.
It was found that the strain on the bottom flange was several times larger than that of the top flange in the previous studies [5,15]. Focusing on these results, the plastic neutral axis of bending in the beams was estimated from strain behavior obtained at section A. The plastic neutral axis at section A of all specimens under cyclic load reversals is shown in Fig. 11(a). The position of the neutral axis can be obtained by using the distance ratio of the height of the beam depth and the strain increment of both bottom and top beam flange. As shown in Fig. 11(a), the plastic neutral axis of SP-1 is located in the vicinity of the center of the beam web, but that of SP-2 is shown to be near the top flange level. Consequently, this caused the strain on the bottom flange to be much larger than that of the top flange (Fig. 11(b)). As shown in Fig. 11(b), while the strain of SP-1 showed a symmetric balance for the top and bottom flange, the test result of SP-2 exhibited that it was concentrated in the bottom flange. The strain of the bottom flange for SP-1 and SP-2 under positive bending is shown in Fig. 11(c). At the same cumulative rotation angle, Rlhl ¼ 0:1 rad, the strain of the composite beam (SP-2) was 4.4 times higher than that of the bare steel beam (SP-1).
5.1.1. Strain concentration and plastic neutral axis The concrete floor slabs may increase the positive flexural capacity, raising the beam neutral axis and, therefore, creating a large strain in the bottom flange.
5.1.2. Moment transfer efficiency of beam web A square tube column had two webs at each side but no web at the center where the beam web was connected. Therefore, the web of the box column was signifi-
5. Discussion 5.1. Composite effects of slabs (Series I—Conventional specimens)
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Fig. 12. Deformation of column flange and strain distribution of connection.
cantly less effective in transferring bending moment due to the out-of-deformation of the column flanges and the loss of the web sections by the weld access hole (Fig. 12(a)). Stress increase at the end of the beam is influenced by the efficiency in transmitting the stress in the web of beam through beam to column joint. Moreover, it was reported that the beam near the connection developed a smaller strength than the beam away from the connection due to the decrease of the web moment transfer as shown in Fig. 12(b) [19]. To investigate the effect of web moment transfer, moment versus curvature relationship was used. The curvature was obtained from the gross sectional average strain of a beam flange. Next, the skeleton curve of a curvature constructed from moment versus curvature relationships and moment versus monotonic curvature relationships at each section under positive bending is shown in Fig. 13. In Fig. 13(a), the moment of section A in bare steel specimen SP-1 appeared 8, 11, and 13% slightly smaller than sections B, C and D at the same curvature, respectively. This means that the web connection did not transfer its full share of the beam moment, resulting in strain concentration of the beam flange. The moment of section A in composite specimen SP-2 exhibited 35, 30, and 36% significantly smaller than sections B, C and D at the same curvature, respectively (Fig. 13(b)). The decrement of the strength of SP-2 was about 3 times of that of SP-1, and this exhibited that composite connection decreased the web moment transfer compared to bare steel connection due to the composite effect of the slab. Therefore, the deformation was concentrated near the beam bottom flange where the constraint was minimal. More detailed studies are required for these results.
5.1.3. Reduction of plastic zone Fig. 14(a) shows the distributions of the curvature in the beam’s longitudinal direction at the completion of the peak moment, and it was assumed that there was almost no curvature change from section D to loading point and it accordingly was simplified into a linear line. Fig. 14(b) shows the ratio of the curvature of each section to the summation of those for all sections. SP-1 had a curvature distribution that changed gradually from section A to section D. But in SP-2, the curvature distribution from section A to section B changed exceedingly and the curvatures from section B to section D were shown to be very small compared with section A. These results also can be seen clearly in Fig. 14(b). Fig. 15 shows the absorbed energy in each section of all specimens. The energy obtained from moment versus curvature hysteresis loops. This plot exhibited that, in the case of SP-2, the ratio of the energy absorbed in section A to that in all sections was near 80%, while, in the case of SP-1, it was a ratio of 59%. Also, while the energy of SP-1 showed a symmetric balance between the positive and negative bending, the test result of SP2 exhibited that it was concentrated in the positive bending direction. These results indicated that, for bare steel connections, the deformation was relatively well distributed over the entire section, but, for composite connections, it had higher local ductility demand in a narrow plasitified zone of the beam. In other words, it can be understood that the plastic zone of composite beams was reduced while bare steel beams secured a wide plastic zone. 5.2. Effect of seismic retrofit (Series II—Retrofit specimens) In previous studies, no weld access hole connections showed outstanding performance in laboratory testing due to the high efficiency of web moment transfer [15,19]. While typical details for new construction used no weld access hole scheme in both the top and bottom web, the retrofit of Specimen SP-3 used it only in the bottom web. In Fig. 13(c), the moment of section A in specimen SP-3 exhibited 18, 8, and 7% smaller than sections B, C, and D, at the same curvature, respectively. The decrement of the strength of SP-3 was significantly smaller than that of SP-2, exhibiting the ratio of 30, 36 and 35%. The decrement of the strength of SP-3 was slightly smaller than that of SP-2. To directly compare moment transfer efficiency of SP-3 with SP-2, moment versus curvature relationship and strain value was plotted in Fig. 16. Fig. 16 indicated that, at the same curvature level, SP-2 and SP-3 specimens had different flexural strength at the critical region, section A. For example, at the curvature of 48 (106/mm), the
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Fig. 13. Moment versus curvature relationships.
strength of the beam for SP-1 already reached 2530 KN.m, while, for SP-2, the strength of the beam was only 2154 KN.m. As shown in Fig. 17, in the same cumulative rotation angle, the strain on the bottom flange of SP-3 was much smaller than that of SP-2 at section A, while it was slightly larger than that of SP-1. Also, Fig. 14 showed that in the case of SP-3, the ratio of the curvature in section A to that in all sections was about 59%, while, in the case of SP-2, it was a ratio of 83%. In addition, the plastic neutral axis of SP-3 was
Fig. 14. Distribution of curvature along beam length.
Fig. 15. Absorbed energy in each section.
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Fig. 16. Effect of no weld access hole scheme.
located slightly lower than that of SP-2 at the same cumulative rotation angle (Fig. 11(a)). These results indicated that the no weld access hole scheme was effective in transferring the web moment even in composite connections, so that it developed less ductility demand on the beam bottom flanges. But, the notch caused by the unsatisfactory workmanship developed and, consequently, SP-3 failed in the bottom flange beneath the weld of the new web plate end prior to developing a sufficient deformation capacity, suggesting that the reliability of the workmanship was needed in doing field constructions of retrofit schemes. The horizontal stiffener specimen (SP-4) improved the ductility of composite connections and showed stable hysteretic characteristics compared to the standard composite specimen (SP-2). This specimen exhibited flange local buckling and reached failure due to a ductile crack initiated at the tip of the horizontal stiffener. Therefore, SP-4 also improved the energy absorption capacity of composite beam connections when compared to specimen SP-2. Fig.13(d) shows that for SP-4, section A and section B of the beam near the column face stiffened by the horizontal stiffener remained in elastic range or deformed only slightly even under ultimate state. The force demand in the existing bottom flange groove weld and the toe of the weld access hole was significantly reduced to a reasonable level, although the plastic neutral axis of this specimen moved forward to the top flange but differed from that of specimen SP-2, as shown in Fig. 11(a). Fig. 14(a) shows the distributions of curvature in the beam’s longitudinal direction at completion of the peak moment. This plot showed that for SP-4, the curvature was small near the connections and increased from the horizontal stiffener end. In addition, in Fig. 14(b), the ratio of the curvatures in sections A, B, C, and D to that in all sections was 7, 8, 50 and 35%, respectively. Similar results can also be seen through the absorbed energy of each section. As shown in Fig. 15, the energy of SP-2 concentrated in section A of 80% compared with all sections. But, SP-4 absorbed energy in section C and section D of 50% compared with all sections. These results may illustrate that moment connections retrofitted using the horizontal stiffener moved the
Fig. 17. Strain distribution.
plastic hinge away from the column face and, thus, achieved a more reliable connection performance even in composite connections. Although the cover plate specimen (SP-5) provides a strengthened zone in the vicinity of the connection that results in a lower level of stress/strain on the weld as shown in Figs.13(e), 14 and 17, the proposed connection scheme exhibited poor performance. As mentioned in the tests results, good workmanship and sound welding quality are still required in order to provide a reliable welded connection. 6. Conclusion The presence of the concrete slabs had an influence on the neutral axis location, the web moment transfer, and the plastic zone of the connection, and these influences caused the strain on the bottom flange to be much larger than that of the top flange. Consequently, the deformation capacity of a composite connection is about half that of a bare steel connection. These results indicated that the detrimental slab effects should be considered in the seismic design of the connection. The no weld access hole detail (SP-3) was effective in the transferring of the web moment even in composite connections, such that it developed less ductility demand on the beam bottom flanges. The cover plate specimen (SP-5) provided a strengthened zone in the vicinity of the connection that can be expected to result in a lower level of stress/strain on the weld. Nevertheless, these proposed connections exhibited poor performance, suggesting that reliability of the workmanship and welding quality were required to ensure the ductile performance of the connection. The horizontal stiffener details (SP-4) showed that an enlarged plastic zone can be achieved and the deformation capacity can be improved. This retrofit scheme demonstrated very good potential in improving the ductility of composite beam connections in existing buildings without removing the floor slabs. Moreover, by using the horizontal stiffener in retrofit of the existing building, workmanship as well as excellent performance
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