- Email: [email protected]
Experimental investigation on the seismic behavior of steel moment connections with decking attachments
Journal of Constructional Steel Research 105 (2015) 174–185
Contents lists available at ScienceDirect
Journal of Constructional Steel Research
Experimental investigation on the seismic behavior of steel moment connections with decking attachments B.W. Toellner a, C.E. Watkins b, E.K. Abbas c, M.R. Eatherton c,⁎ a b c
Thornton Tomasetti, 2000 L St. NW, Suite 800, Washington, DC 20036, USA Walter P. Moore, 1301 Mckinney, Suite 1100, Houston, TX 77010, USA Department of Civil and Environmental Engineering, Virginia Tech, 200 Patton Hall, Blacksburg, VA 24061, USA
a r t i c l e
i n f o
Article history: Received 25 June 2014 Accepted 15 November 2014 Available online 15 December 2014 Keywords: Steel moment resisting frames Protected zone Powder actuated fasteners Puddle welds Low cycle fatigue fracture Seismic behavior
a b s t r a c t Steel moment resisting frames rely on large inelastic strains in the beam plastic hinge region to dissipate seismic energy during an earthquake and protect the building against collapse. To limit the potential for premature fracture and because of a lack of test data, fasteners, attachments and defects are prohibited in the plastic hinge region, also referred to as the protected zone in the AISC Seismic Provisions. However, unauthorized attachments and defects occur in many buildings in practice. A set of twelve full-scale moment connection tests were conducted to explore the effect of powder actuated fasteners (PAFs) and puddle welds on the seismic performance of steel moment connections. Both reduced beam section and extended end plate connections were tested with W24 × 62 and W36 × 150 beams. Five specimens included PAFs or puddle welds representing typical steel deck attachment to the top flange of the beam. Three of the specimens included PAFs in a grid over the top and bottom flange and on the web. All twelve specimens passed the qualification criteria for special moment resisting frames (SMRFs) in the AISC Seismic Provisions as they were subjected to a cyclic displacement protocol up to 4% story drift while retaining 80% of their nominal plastic moment capacity. Therefore, the tested moment connection configurations with PAFs and puddle welds were found to produce ductile SMRF type seismic performance. Furthermore, PAFs and puddle welds were found to have negligible effect on cyclic envelope, moment capacity, energy dissipation and strength degradation prior to fracture. © 2014 Elsevier Ltd. All rights reserved.
1. Introduction Special moment resisting frames (SMRFs) depend on large inelastic strains in the beam-to-column connections to dissipate seismic energy and protect buildings from collapse. The ANSI/AISC 341-10 Seismic Provisions for Structural Steel Buildings [2] define the ends of the beams in a SMRF as protected zones at the locations where large inelastic strains are expected. The term “protected” refers to the concept that this portion of the beam should be protected from defects, fasteners, and other discontinuities because their effect on low cycle fatigue fracture is not well understood. For example, AISC 341 [2] prohibits any decking attachments that penetrate the beam flange and any welded, bolted, screwed, or shot-in attachments for other purposes. The extents of the protected zone are defined for each prequalified moment connection included in AISC 358 [3] and are shown graphically in Fig. 1 for two types of SMRF connections. The concept of the protected zone and related restrictions is largely based on a lack of data. In practice, however, the restriction on fasteners in the protected zone creates issues in construction due to the difficulty in communicating and enforcing the requirements. Many instances of ⁎ Corresponding author. E-mail address: [email protected] (M.R. Eatherton).
http://dx.doi.org/10.1016/j.jcsr.2014.11.006 0143-974X/© 2014 Elsevier Ltd. All rights reserved.
unauthorized attachments are submitted to engineers of record each year, and for each reported case of unauthorized attachment, there are likely many more fasteners and defects in the protected zone that go unreported. However, since moment connection tests in the literature focus on the behavior of the connection itself and thus have wellcontrolled defect free plastic hinge regions, there is almost no data on cyclic moment connection behavior with defects or fasteners other than welded shear studs. As a result, engineers of record are faced with a difficult decision between leaving the unauthorized attachment, or requiring an invasive repair method without having any data as to the expected seismic behavior associated with either option. A series of twelve full-scale beam-to-column moment connection tests were conducted to evaluate the effects of powder actuated fasteners (PAFs) and puddle welds applied in the protected zone on the seismic behavior of steel moment connections. The specimens are intended to represent common types of attachments made to the beam in the protected zone such as attachment of steel deck to the beam top flange, attachment of cold-formed steel wall track to the underside of the beam, and attachments for mechanical/electrical nonstructural systems. The objective of the testing program was to investigate whether powder actuated fasteners or puddle welds affected the moment capacity, strength degradation, energy dissipation, and fracture potential of the moment connections as compared to
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
175
To the end of the RBS
Shaded Area is the Protected Zone
RBS Cut a) Reduced Beam Section Moment Connection
b) After Installation
a) Distance equal to the smaller of d or 3bf
Fig. 2. Hilti X-ENP powder actuated fastener.
Flange Width, bf
d Shaded Area is the Protected Zone
b) Unstiffened End Plate Moment Connection Fig. 1. Examples of protected zone for two moment connection types.
specimens with no fasteners. The special moment resisting frame (SMRF) qualification criteria of AISC 341-10 [2] were used as a benchmark for defining acceptance as the specimens were considered to produce sufficiently ductile behavior if they could retain a significant portion of their moment capacity after being subjected to cyclic loading up to 4% story drift. 2. Background There have been many experimental programs on full-scale moment connections since the 1994 Northridge Earthquake and the 1995 Kobe Earthquake exposed unexpected fractures in steel beam to column connections. Experimental programs focused on a range of topics including reduced beam section connections (e.g. [6]), extended end plate connections (e.g. [1]), effect of composite slabs (e.g. [9]), panel zone strength (e.g. [11]), and much more. A subset of these testing programs included a composite slab with welded shear studs attached to the beam in the protected zone (e.g. [5,7,9,12,14,18]). Out of this body of tests, one beam-to-column connection specimen with welded shear studs in the protected zone experienced fracture of the flange initiating at the shear stud [14]. It was concluded that the fracture of the beam flange was a direct result of the reduced notch toughness of the base metal caused by shear stud welding. The other specimens in this testing program and specimens from other testing programs did not develop fracture at the welded stud location. Besides the effect of the fastener itself on stress concentrations and material properties, connections between the concrete slab and the steel beam strongly affect the strain distribution in the beam. Beamto-column tests using composite slabs and welded shear studs through the protected zone have demonstrated that composite action in the connection region causes significantly increased strain demands in the beam bottom flange [9,12,17]. Both the rapid change from composite section to bare steel beam right at the connection to the column and larger strain demands in the critical bottom flange where the majority of fractures were observed after the Northridge earthquake were considered detrimental.
However, the effect of a particular type of fastener on stress concentrations, base metal properties, and degree of composite action is heavily dependent on the type and configuration of fasteners. For this reason, it is necessary to investigate the effect of each type of artifact (i.e. fastener or defect) on moment connection seismic behavior separately. Powder actuated fasteners (PAFs) are a common method for attaching steel deck to the top flange of steel beams, as well as a common means for attaching nonstructural elements such as cold-formed steel tracks and mechanical, electrical, and plumbing elements to the structure. The fastener (example shown in Fig. 2) is driven into the steel plies using a fastening tool. The effect of PAFs on fracture of steel coupons subjected to monotonic tension has been studied by Beck and Engelhardt [4] and it was determined that coupons with PAFs had higher tensile strength than coupons with equivalent drilled holes. Moreover, a typical coupon specimen with PAF was shown to reach a strain of 17% before fracture which is less than a specimen with no holes, but approximately 70% more ductility than a similar specimen with drilled holes. The increased strength and ductility of steel coupons with PAFs as compared to drilled holes might be attributed to increased strength in the surrounding base metal caused during the fastener application or to residual compressive stresses in the material surrounding the hole. Studies have also been conducted to investigate the high cycle fatigue performance of steel with PAFs (e.g. [13]), tube connections using PAF (e.g. [10]), and steel deck attachment using PAF (e.g. [15]). The extent of material affected by PAF was investigated [20] in which it was found that a particular type of PAF affected the Vickers hardness up to 6 mm from the edge of the fastener. However, the behavior of steel with PAFs subjected to cycles of large inelastic strains, such as those experienced in moment frame plastic hinges, has not been experimentally investigated. 3. Testing program description 3.1. Test setup The test setup is shown in Figs. 3 and 4. The configuration uses a W14 × 257 vertical column, approximately 3.67 m tall that is restrained against lateral translation at the top and bottom. The W24 × 62 and W36 × 150 beam specimens attach to the column with a bolted end plate connection to allow the same column to be used for every test. Furthermore, the two ends of the beam were used for two different tests to maximize the efficiency of steel used for these tests. Additional details on geometry and reaction frames are given in Toellner [16] and Watkins [19]. Load was applied with an MTS model 201.70 actuator with force capacity of 956 kN in tension and 1468 kN in compression and total stroke equal to 508 mm. The geometry of the test setup is intended to simulate an exterior column subassemblage of a moment frame undergoing rotation due to
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
W14x257 Column
3670 mm
176
MTS servo-controlled actuator: 965 kN capacity and 508 mm stroke P= Actuator Force
Lateral bracing locations
Beam CL
Table 1 Cyclic displacement protocol. Percent story drift (%)
Number of cycles
Displacement at actuator (mm)
0.38 0.50 0.75 1.00 1.50 2.00 3.00 4.00 4.66
6 6 6 4 2 2 2 2 Up to 5
19 26 39 52 78 104 155 207 241
1543mm for W24x62 1359 mm for W36x150
Beam rotated and flipped for subsequent specimen Lf = 4973 mm L c = 5182 mm
Fig. 3. Test frame layout.
story drift. The 5.18 m distance from the column centerline to the actuator represents the distance from the column centerline to the beam inflection point in a 10.36 m long bay assuming an idealized moment diagram. Similarly, the column height of 3.67 m represents the distance between column inflection points on adjacent floors assuming an idealized moment distribution and implying a building with 3.67 m floor heights. A third reaction frame provided lateral restraint to the test specimens. The location of the frame could be moved depending on unbraced length requirements of the specimens. The resulting lateral bracing points were 0.84 m, 1.49 m, and 4.65 m from the face off the column for W24 × 62 beam specimens and 1.14 m, 1.79 m, and 4.65 m for the W36 × 150 beam specimens. 3.2. Loading protocol and qualification criteria The specimens were subjected to the cyclic story drift history described in AISC 341-10 for special moment frames [2] and given in Table 1. The target displacements at the actuator location were determined by multiplying the target story drifts by the distance between the actuator and the column centerline, Lc = 5.18 m. Beyond the qualification amplitude of 4% story drift, additional cycles were performed at an amplitude of 4.7% story drift which corresponds to the maximum stroke of the actuator. The tests were displacement controlled using
Fig. 4. Photograph of test setup.
external displacement feedback from a string potentiometer located below the actuator with a displacement rate of 61 mm/min for all tests. The SMRF qualification criteria from AISC 341-10 [2] were used as a benchmark to determine whether the connections containing PAFs or puddle welds achieved sufficient ductility to produce SMRF type of performance with the desirable seismic behavior implied therein. The SMRF qualification criteria in AISC 341-10 chapter K specify that a specimen must undergo one complete cycle at 4% story drift while sustaining a moment resistance at the face of the column that is at least 80% of the nominal plastic moment capacity. The nominal plastic moment capacity is computed with nominal yield stress and nominal plastic section modulus.
3.3. Test specimens The test program consisted of twelve specimens as listed in Table 2. This set of specimens includes variation in beam depth (nominal W24 to W36), flange thickness (15 mm to 24 mm), flange width (178 mm to 305 mm), and beam weight (92 kg/m and 223 kg/m). The flange thickness may be an important variable as the location of the PAF tip in the depth of the flange may affect the potential for fracture initiation and both the flange thickness and width will affect the relative magnitude of the disturbance in the strain field. The depth of the section will affect the inelastic strain demands as the distance from the neutral axis to the fastener will vary for the same curvature demand. The flange width and incorporation of RBS cuts will further affect inelastic strain demands at the flange tips and thus are also key variables that were varied in this study. Six of the specimens included a reduced beam section (RBS), but also utilized an extended end plate so that the same column could be reused. It is noted that while RBS specimens don't typically include endplates, it is expected that the relatively rigid endplate did not have a significant effect on the inelastic deformations of the RBS plastic hinge which are the focus of this investigation. The end plate connections were designed in accordance with ANSI/AISC 358-10 for the full strength of the beam without RBS and additional details can be found in Eatherton et al. [8]. A similar variation of fasteners was used for both beam depths as shown in Fig. 5. Fasteners at 305 mm spacing are typical for deck attachments to the top flange or partition attachments to the underside of the bottom flange so both PAF and puddle weld configurations with this spacing were included in the testing program. Three specimens (Specimens 6, 9, and 12) included a grid of PAFs on both flanges with spacing equal to 51 mm, equally spaced lines of PAF across the width of the flange with spacing not less than 51 mm, and distance from center of PAF to the edge of the flange of 25 mm. Both Specimen 9 and Specimen 12 also incorporated a grid of PAFs on one side of the web as depicted in Fig. 5m. Although the specimens with a grid of PAF don't represent attachments typically made in practice, the grid specimens were intended to identify the worst case locations as PAF could conceivably be located in any of these positions. Assuming negligible interaction between the fasteners, a single PAF grid test can therefore represent a multitude of individual tests in which a PAF is present at any of the given locations.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
177
Table 2 Matrix of test specimens. Specimen
Beam size
1 2 5b 3b 4 6 7 8 9 10 11 12
W24 × 62 W24 × 62 W24 × 62 W24 × 62 W24 × 62 W24 × 62 W36 × 150 W36 × 150 W36 × 150 W36 × 150 W36 × 150 W36 × 150
Connection type
Protected zone lengtha (mm)
Nominal Mp and 0.8Mp (kn m)
RBS RBS RBS BUEEP BUEEP BUEEP RBS RBS RBS BSEEP BSEEP BSEEP
620 620 572 572 620 572 1003 1003 1003 818 818 818
556/445 556/445 556/445 865/691 865/691 865/691 2261/1810 2261/1810 2261/1810 3281/2626 3281/2626 3281/2626
Fasteners in the protected zone None 4 PAFs spaced at 305 mm 4 puddle welds spaced at 305 mm None 4 PAFs spaced at 305 mm PAF grid None 4 puddle welds spaced at 305 mm PAF grid None 4 PAFs spaced at 305 mm PAF grid
RBS = reduced beam section. BUEEP = bolted unstiffened extended end plate. BSEEP = bolted stiffened extended end plate. a Protected zone length measured from column face. b Specimen 3 conducted prior to Specimen 5, but reordered for test matrix.
fastener size was chosen to represent the largest size PAF used in deck attachment to steel beams. The arc spot puddle welds were applied by a qualified lab technician who is certified for this type of welding and other types of welding by the American Welding Society. The puddle welds were 19 mm in diameter and made through a small square of 20 gage metal decking and then the excess decking material was trimmed off using a torch to allow strain gage application.
It is noted however, that the grid of PAF on one specimen is an extreme condition that is not indicative of construction practice. 3.4. Power actuated fasteners and puddle weld information The PAFs used in this study were Hilti X-ENP-19 L15 nails driven into test specimens with a 0.27 caliber Hilti DX 76 powder-actuated tool. The
762
127
458
203
2 Spaces at 304
247
337
No Fasteners No Fasteners
a) Specimen 1 54
3 Spaces at 304
2 Spaces at 304 Powder Actuated Fastener (PAF)
PAF
g) Specimen 7 247
k) Specimen 11 337
229
51 Typical
b) Specimen 2
No Fasteners Puddle Welds
c) Specimen 3
l) Specimen 12
54
3 Spaces at 304
51 Typical
PAF
889
h) Specimen 8
229
127
PAF
d) Specimen 4 54
3 Spaces at 304
PAF
991
Puddle Welds
PAF
152 Typical
i) Specimen 9
e) Specimen 5
889
51
51 Typical
No Fasteners
m) PAF Applied to One Side of
711
f) Specimen 6
j) Specimen 10 PAF Fig. 5. Specimen configurations.
Web for Specimens 9 and 12
178
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
4. Test results
Table 3 Measured steel properties from tension coupons.
W24 × 62
W36 × 150
Yield stress, Fy (MPa)
Ultimate stress, Fu (MPa)
Elongation at fracture using punch marks (%)
363 365 360 363 364 374 376 372
489 478 461 476 479 470 478 476
22 27 23 24 30 30 29 30
Top flange Bottom flange Web Average Top flange Bottom flange Web Average
3.5. Ancillary material tests All beam specimens were specified to be ASTM A992 structural steel with a minimum yield stress of 345 MPa and a minimum tensile strength of 448 MPa. All beams of the same size were obtained from the same heat of material. Tension tests were performed on six coupon specimens cut with longitudinal orientation from the top flange, bottom flange, and web from the undeformed regions of one W24 × 62 beam and one W36 × 150 beam in accordance with ASTM A6/A6M-12 [21]. The geometry of the coupon specimens was in accordance with ASTM 370-07a [22] with a width of 38 mm, a total length of 457 mm and a gage length of 203 mm. Each specimen included punch marks at the ends of the gage length for which the distance between punch marks was measured before and after testing. The yield stress, ultimate stress, and elongation at fracture as measured using the distance between punch marks are given in Table 3.
All twelve specimens tested in this experimental program satisfied the SMRF qualification criteria of AISC 341-10. This means that moment connections with configurations of fasteners such as shown in Fig. 5 can produce SMRF type of behavior with all the ductility and energy dissipation implied therein. The results of each group of tests are presented in this section, and then the results are compared and synthesized in the following section. The plots of load-deformation presented in this section show the moment at the face of the column given by the applied force, P, multiplied by the distance to the face of the column, Lf (see Fig. 3) and normalized by the nominal plastic moment capacity, Mp, calculated using nominal section properties and nominal yield stress. The horizontal axes are story drift calculated as the displacement at the actuator, δSP2 (see Fig. 6) divided by the distance to the column centerline, Lc (see Fig. 3).
4.1. Behavior of W24 × 62 specimens with RBS The progression of limit states was similar for all specimens in this group starting with significant yielding in the extreme fibers at the reduced section, spread of plasticity, local buckling of the flanges in association with out-of-plane buckling of the web, crack initiation typically at the flange tips on the inside face of the local buckle, and fracture propagation through an entire flange. All specimens experienced some amount of fracturing at the minimum flange section. In all cases, the fractures initiated at the edge of the flange on the inside of a local buckle and propagated toward the center of the beam flange. Fig. 7 shows the
Load Cell and Displacement Sensor in the Actuator
LVDT 8 Inclinometer 1 Inclinometer 2 Inclinometer 3
LVDT 1 LVDT 5 LVDT 6 LVDT 2
LVDT 4
Story Drift, δ SP2 / Lc (%)
Flange Fracture
LVDT 9 LVDT 7
String Potentiometer 3
LVDT 3
Flange Fracture
a) Specimen 1, No Fasteners
Applied Moment , PLf / Mp
The instrumentation plan was designed to include enough displacement sensors to decompose the applied story drift into components due to the reaction frame, column outside the panel zone, column panel zone, end plate, plastic hinge, and elastic deformations of the beam outside the plastic hinge region. Nine LVDTs, three string potentiometers, and three inclinometers were used as shown in Fig. 6. Additional details are included in Eatherton et al. [8]. In addition to displacement and rotation sensors, a set of high strain rated strain gages were applied in configurations across the flange width with the intent to investigate differences in the strain gradient with and without fasteners. The strain gages had a published strain range of 10% to 15% and were applied in a line of seven strain gages on the outside of the flange and four on the inside face of the flange. In locations with a fastener at the center of the flange, only six strain gages were applied on the outside of the flange.
Applied Moment, PLf / Mp
3.6. Instrumentation
String Potentiometer 2 δSP2
Note: LVDT = Linear Variable Differential Transformer
Fig. 6. Instrumentation plan.
Story Drift , δ SP2 / Lc (%)
b) Specimen 2, PAF at 305 mm Fig. 7. Load-deformation behavior of W24 × 62 specimens with RBS.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
load-deformation behavior for Specimen 1 without any fasteners, and Specimen 2 with PAF at 305 mm spacing. Both specimens experienced fracture during the second cycle at 4.7% story drift and Specimen 5 (not shown) experienced fracture in the third cycle at 4.7% story drift. The strength degradation was slightly delayed for Specimen 2, but this was likely due to slight differences in the specimen geometry and material. The load-deformation response for Specimen 5 was similar and thus not shown. The fracture surfaces were analyzed to determine the crack initiation locations and fracture propagation paths. Fig. 8c and d shows that the fracture in Specimen 1 with no fasteners started at the flange tips, with initiation occurring at the inside of the local buckles. Crack initiation at the flange tips was also noted in Specimen 2 in which the fracture propagated toward the center of the section displaying five distinct cycles of ductile fracture propagation before the crack reached the PAF location and brittle overload occurred (see Fig. 8a). Although a similar process occurred in Specimen 5 including crack initiation, limited ductile fracture propagation, and brittle overload as shown in Fig. 8b, there was less ductile fracture propagation than Specimen 2. It is noted that the fracture in Specimen 5 occurred in the unadulterated flange opposite from the one that included puddle welds. 4.2. Behavior of W24 × 62 specimens without RBS The progression of limit states was similar for all specimens in this group starting with significant yielding in the extreme fibers near the endplate connection, spread of plasticity, and local buckling of the flanges in association with out-of-plane buckling of the web. None of the specimens experienced significant fractures even after being subjected to multiple cycles at 4.7% story drift.
Ductile Overload
Cycle Marks
PAF Location
179
Fig. 9 shows relatively similar load-deformation behavior between Specimen 3 which had no fasteners and Specimen 6 which had a grid of PAF on both flanges as shown in Fig. 5f. Specimen 4 with PAF at 305 mm exhibited behavior similar to Specimen 3 and is thus not discussed further here. See Eatherton et al. [8] for more information. Testing of Specimen 3 (as well as Specimen 4 which is not shown) was arbitrarily stopped after 2 cycles at 4.7% story drift, but no fractures were observed. Testing of Specimen 6 was continued for five cycles at 4.7% story drift until such time as the load-deformation behavior appeared to be stabilizing, implying that the inelastic local buckling (see Fig. 10) was not varying much from one cycle to the next. At a few of the PAF locations at the peaks of the local buckles, the hole elongated and small tears formed radiating out from the PAF (see Fig. 10). As compared to the cracks that initiated at the flange tips of the RBS in the previous group of specimens, these tears appeared to propagate very slowly, extending approximately 10 mm from the edge of the PAF by the end of the test. 4.3. Behavior of W36 × 150 specimens with RBS The W36 × 150 specimens experienced a similar progression of limit states as the W24 × 62 specimens starting with yielding of the extreme fibers at the reduced section, spread of plasticity, local buckling of the flange in association with out-of-plane buckling of the web, crack initiation, and fracture propagation. However, unlike the W24 × 62 specimens, the cracks initiated on the face of the flanges at the inside of the local buckles as shown in Fig. 11a. For Specimen 7, the cracks were identified during the 4% story drift cycles and then during the 4.7% story drift cycles, fracture propagated through the flange and web as shown in Fig. 11c. A similar process occurred for Specimen 9 that had a grid of
Critical Flow Region 2 1
Crack Initiation and Slow Growth
Final Overload Region
Crack Initiation
Shear Lips
Brittle Overload
a) Specimen 2 Failure Origin
c) Specimen 1
Crack Initiation Shear Lips
b) Specimen 5 Internal Cracks
d) Specimen 1
Fig. 8. Analysis of fracture surfaces for W24 × 62 specimens with RBS.
Failure Origin
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
Applied Moment , PL f / Mp
180
cycle at 4% story drift. Specimen 8, which is not shown here, experienced a fracture at the toe of the stiffener which was unrelated to the puddle welds that were present on that specimen. As shown in Fig. 12, the load-deformation behavior of the specimen without fasteners, and the specimen with a grid of PAF on both flanges was quite similar prior to fracture. The differences in the peak strength and strength degradation will be analyzed further in a subsequent section.
Test Stopped Without Fracture
4.4. Behavior of W36 × 150 specimens without RBS
Story Drift , δ SP2 / Lc (%)
Applied Moment , PL f / M p
a) Specimen 3, No Fasteners Test Stopped Without Fracture
The progression of limit states for the W36 × 150 specimens that didn't include an RBS was the same as those for the W24 × 62 specimens with no RBS. Although some tearing was observed at the stiffener toe, Specimens 10 and 11 with no fasteners and PAF at 305 mm spacing respectively, did not experience significant fracture as they were subjected to the full loading protocol including five cycles at 4.7% story drift. Specimen 12 that included a grid of PAF over both flanges and web, experienced a fracture during the fourth cycle at 4.7% story drift. However, the load-deformation behavior of Specimens 10 and 12 was quite similar as shown in Fig. 13. Specimen 12 underwent substantial local buckling and inelastic strains prior to fracture as shown in Fig. 14a. The fracture in specimen 12, shown in Fig. 14b, initiated on the surface of the flange at a PAF located on the inside of a local buckle. 5. Discussion of results
Story Drift, δ SP2 / L c (%)
b) Specimen 6, Grid of PAF
This section summarizes several studies performed to compare test data in order to identify differences between the performance of control specimens and those with PAFs or puddle welds. These studies include comparisons of hysteretic behavior, strength degradation and energy dissipation.
Fig. 9. Load-deformation behavior of W24 × 62 specimens without RBS.
5.1. Fracture behavior PAF over both flanges. As shown in Fig. 11b, the fracture occurred at the reduced section where the local buckle had the largest curvature. An analysis of the fracture surface, shown in Fig. 11d, revealed that the crack initiated at the surface of the flange around a PAF at the inside of the local buckle. These cracks were observed to form during the second
Tear
Fig. 10. Specimen 6 at the end of the test.
Several general observations are made regarding the fracture behavior of the twelve specimens. Table 4 provides a summary of the fractures and shows that all six RBS specimens experienced some degree of fracture during 4.7% story drift cycling. Specimen 1 experienced a ductile flange tear and the other five RBS specimens experienced brittle fractures along an entire flange width and a portion of the web. The only non-RBS specimen to experience a fracture was Specimen 12 which contained a grid of PAF over both flanges and the web. Strain gage data (not shown here, see Eatherton et al. [8]) did not reveal noticeable differences in flange strain gradient between specimens with PAF, puddle welds or no fasteners. At large story drifts, inelastic strains were found to be dominated by curvature related to local buckles and inelastic strain was not found to concentrate in the vicinity of PAF or puddle welds. For the W24 × 62 beam specimens with RBS, the fractures occurred during the 2nd or 3rd cycle at 4.7% story drift regardless of whether there were PAF, puddle welds, or no fasteners included in the protected zone. This suggests that although the PAF may interact with the fracture process, the configurations of PAF and puddle welds tested did not cause fracture significantly earlier than specimens with no fasteners. The W36 × 150 RBS specimen with no fasteners fractured during the 5th cycle at 4.7% story drift, whereas the specimens with a grid of PAF and puddle welds fractured in the 2nd and 1st cycles at 4.7% story drift respectively. However the fracture in the specimen with puddle welds (Specimen 8) was not related to the puddle welds, but instead occurred at the junction of the stiffener and flange. This implies some inherent variability in the cycle at which fracture occurs. Specimen 12, which included a grid of PAF, fractured during the 4th cycle at 4.7% story drift whereas the other non-RBS specimens (Specimens 10 and
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
a) Specimen 7
181
b) Specimen 9 Crack Initiation
Brittle Overload Propagation Direction
Critical Flow Region Shear Lips
c) Specimen 7 Fracture Surface Crack Initiation
Powder Actuated Fastener
Brittle Overload Ductile Overload and Critical Flow Region
Shear Lips
d) Specimen 9 Fracture Surface Fig. 11. Fracture for Specimen 7 without fasteners and Specimen 9 with grid of PAF.
11) did not experience fracture through five cycles at 4.7% story drift even though Specimen 11 included PAF at 305 mm spacing. The results for the W36 × 150 specimens that included a grid of fasteners suggest that one or more PAF applied in critical locations such as the inside of a local buckle may lead to fracture during an earlier cycle than specimens without fasteners. The potential for fracture is supported by the small tears that formed at PAF in critical locations after a large number of inelastic strain cycles. It is noted, however, that fracture occurred after SMRF qualification had been satisfied, the difference in the cycle number at fracture was small, and that the two W36 × 150 PAF specimens that fractured included a grid of PAF that represents an extreme condition, not indicative of construction practice. Since the sample size was small and the difference in the cycle number at fracture
was small, it was not possible to evaluate the difference in a statistically significant manner.
5.2. Load-deformation behavior Moment-drift envelopes for all tests are shown in Fig. 15. The envelope was developed by extracting the peak moments for any cycle at each story drift level. It is shown in Fig. 15a that the two groups of W24 × 62 specimens had nearly identical envelopes with slight variations that were likely not related to the puddle welds or PAF. Similarly, Fig. 15b shows that the cyclic envelopes for the W36 × 150 specimens were not affected by inclusion of puddle welds or PAF. It is expected
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
Applied Moment, PL f / M p
Flange Fracture
Applied Moment, PLf / Mp
182
Test Stopped Without Fracture
Story Drift , δ SP2 / Lc (%)
Story Drift , δ SP2 / Lc (%)
Flange Fracture
a) Specimen 10 without Fasteners
Applied Moment, PLf / Mp
Applied Moment , PL f / Mp
a) Specimen 7 without Fasteners
Flange Fracture
Story Drift , δ SP2 / Lc (%)
b) Specimen 9 with Grid of PAF Fig. 12. Load-deformation behavior of W36 × 150 specimens with RBS.
that the accelerated strength degradation of Specimen 11 was due to larger initial imperfections, not the PAF. The similarity between the envelopes with and without fasteners in the protected zone is also demonstrated in Table 5. The maximum moment capacities are all within 1% and 5% of the average value for their group for W24 × 62 and W36 × 150 specimens, respectively. Similarly, the moment capacities at the qualification cycle were within 5% of their group average for all groups except the W36 × 150 specimens with no RBS which experienced more variability with as much as 12% difference from the average moment capacity. Table 5 also shows that all twelve specimens satisfied the AISC 34110 qualification criteria for special moment resisting frames by retaining 80% of the nominal moment capacity, Mp, through one full cycle at 4% story drift. Furthermore, the moment capacity during the 4% story drift qualification cycle was on average 37% greater than the 0.8Mp.
Story Drift , δ SP2 / Lc (%)
b) Specimen 12 with Grid of PAF Fig. 13. Load-deformation behavior of W36 × 150 specimens without RBS.
the non-RBS specimens showed strength degradation ratios that increased as compared to the 4% cycles implying that local buckling deformations began to stabilize. 5.4. Energy dissipation The energy dissipation behavior of each specimen was also compared for each group of specimens. Energy dissipation was calculated using the trapezoidal rule for numerical integration and the resulting values for specimens with RBS are shown in Fig. 17. At the conclusion of the 4% story drift cycles, the energy dissipation for the specimens with PAF or puddle welds was on average within 3% of the energy dissipated by the associated specimens without fasteners. The energy dissipation at the end of the test varied more as it was highly dependent on the number of cycles performed at 4.7% story drift level.
5.3. Strength degradation 6. Summary and conclusions A strength degradation ratio, η, was defined for a given story drift amplitude, i, as the ratio of the moment at the second peak displacement to the moment at the first peak displacement. Despite its designation as a strength degradation ratio, values greater than 100% are possible and indicate an increase in moment capacity due to strain hardening. Fig. 16 shows the strength degradation ratio for the specimens without RBS. It is shown that within each group, the specimens with and without fasteners exhibit very similar strength degradation trends. For RBS specimens (not shown here), strain hardening controls through the 2% story drift cycles, followed by strength degradation associated with local buckling. Specimens without RBS showed strain hardening through the 3% story drift cycles suggesting that local buckling was delayed as compared to RBS specimens. During the 4.7% story drift cycles,
Twelve full-scale beam-to-column moment connection tests were conducted with a range of variables to investigate the effect of powder actuated fasteners and puddle welds on the seismic behavior of moment frames. Variations in the test specimens included control specimens with no fasteners, PAF at 305 mm spacing representing typical deck attachment, puddle welds at 305 mm spacing, and PAF in a dense grid over the plastic hinge region with 25 mm spacing to any edge. Sections tested included W24 × 62 and W36 × 150 beams with flange thicknesses of 15 mm and 24 mm, respectively. Reduced beam section (RBS) connections and non-RBS connections were considered representative of the range of currently prequalified special moment resisting frame connection types.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
183
1000 Test Test Test Test Test Test
800
Applied Moment (kN-m)
600 400
1: 2: 5: 3: 4: 6:
RBS24 RBS24-PAF12 RBS24-PW12 W24 W24-PAF12 W24-PAF-array
200 0 -200 -400 -600 -800 -1000 -5
-4
-3
-2
-1
0
1
2
3
4
5
3
4
5
Story Drift (%)
a) W24x62 Envelopes 5000 Test Test Test Test Test Test
4000
Crack Initiation
Shear Lips
3000
Applied Moment (kN-m)
a) Specimen 12 Buckled Shape
2000
7: RBS36 8: RBS36-PW12 9: RBS36-PAF-array 10: W36 11: W36-PAF12 12: W36-PAF-array
1000 0 -1000 -2000 -3000
Brittle Overload
-4000 -5000 -5
-4
-3
-2
-1
0
1
2
Story Drift (%)
Powder Actuated Fasteners
b) W36x150 Envelopes Fig. 15. Moment vs. story drift envelopes extracted from cyclic test data.
b) Specimen 12 Fracture Surface Fig. 14. Behavior of Specimen 12 with grid of PAF over both flanges and the web.
All twelve specimens satisfied the SMRF qualification criteria in ANSI/AISC 341-10 Chapter K, possessed significant ductility, demonstrated substantial energy dissipation, and according to ANSI/AISC 3410-10 would be considered adequate for use in a special moment resisting frame. The behavior of the twelve specimens was quantitatively compared using several different approaches. The cyclic-load deformation behavior and cyclic envelopes were compared for specimens with no fasteners, puddle welds, and PAF. It was found that the fasteners
had negligible effect on the hysteretic behavior or envelope prior to fracture. Furthermore the achieved moment capacities during the qualification cycle and maximum moment capacities were unchanged with the inclusion of puddle welds and PAF. The effect of puddle welds and PAF on strength degradation and energy dissipation was similarly found to be negligible prior to fracture, even in the specimens with a dense grid of PAF. Although the W24 × 62 specimens suggest that the PAF and puddle welds did not change which cycle saw fracture, the results for the two W36 × 150 specimens that included a grid of fasteners suggest that
Table 4 Summary of fracture cycle and location. Specimen number
Specimen description
Cycle at significant fracture
Fracture location
1 2 5 3 4 6 7 8 9 10 11 12
W24 W24 W24 W24 W24 W24 W36 W36 W36 W36 W36 W36
2nd cycle at 4.7% story drift 2nd cycle at 4.7% story drift 3rd cycle at 4.7% story drift N/A N/A N/A 5th cycle at 4.7% story drift 2nd cycle at 4.7% story drift 1st cycle at 4.7% story drift N/A N/A During 4th cycle at 4.7% story drift
Bottom flange, near center of RBS Through PAF location at center of RBS on top flange Bottom flange, away from puddle welds, near center of RBS Test stopped after two cycles at 4.7% No fracture Test stopped after two cycles at 4.7% No fracture Test stopped after five cycles at 4.7% No fracture Top flange at center of RBS, initial fracture partial depth then opened Bottom flange at tip of end-plate stiffener, not at puddle welds Top flange near center of RBS Through two PAFs Test stopped after five cycles at 4.7% No fracture Test stopped after five cycles at 4.7% No fracture Bottom flange through multiple PAF and up to a PAF located on the web
× × × × × × × × × × × ×
62 with RBS but no fasteners 62 with RBS and PAF at 305 62 with RBS and puddle welds at 305 62 without RBS or fasteners 62 without RBS and PAF at 305 62 without RBS and grid of PAF 150 with RBS and no fasteners 150 with RBS and puddle welds at 305 150 with RBS and grid of PAF 150 without RBS and no fasteners 150 without RBS and PAF at 305 150 without RBS and grid of PAF
184
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
W24 × 62
RBS
1 2
No RBS
W36 × 150
RBS
No RBS
Fasteners
None PAF at 305 mm Puddle welds None PAF at 305 mm Grid of PAF None Puddle welds Grid of PAF None PAF at 305 mm Grid of PAF
5 3 4 6 7 8 9 10 11 12
Max. moment capacity (kN m)
Moment capacity 4% drift (kN m)
710 705
545 586
701 1007 994
548 923 933
998 3196 2943 3045 4131 3856
961 2835 2756 2698 3734 3091
4108
3638
Strength Degredation Ratio (%)
250 200 150 100 50 0
1
2
3
4
5
a) W24x62 with RBS 2500 Test 7: RBS36 Test 8: RBS36-PW12 Test 9: RBS36-PAF-array
2000
1500
1000
500
0
0
1
2
3
4
5
Story Drift (%)
b) W36x150 with RBS Fig. 17. Energy dissipation for each group of specimens.
with no fasteners through the SMRF qualification cycles. Since all twelve specimens satisfied the qualification criteria, it has been demonstrated that moment connections with puddle welds and PAF in configurations such as those tested in this study are capable of substantial ductility consistent with SMRF behavior.
110 Test 3: W24 Test 4: W24-PAF12 Test 6: W24-PAF-array
100 95
Acknowledgments
90
This material is based on work supported by the Hilti, Inc. and the American Institute of Steel Construction. In-kind funding was provided by the Banker Steel and Applied Bolting Technology.
85 80 -5
-4
-3
-2
-1
0
1
2
3
4
5
Story Drift Ratio (%)
References
a) W24x62 without RBS Strength Degredation Ratio (%)
300
Story Drift (%)
one or more PAF applied in critical locations (i.e. the inside of a local buckle) may lead to fracture during an earlier cycle in the displacement protocol. In the three specimens with grids of fasteners (Specimens 6, 9 and 12), tears initiated at PAF in critical locations. It is noted, however, that fracture occurred after SMRF qualification had been satisfied, the difference in the cycle number at fracture was small, and that the two W36 × 150 specimens that fractured included a grid of PAF that represents an extreme condition, not typical of construction practice. On the other hand, PAF or puddle welds applied in a pattern typical of deck attachment appeared to have no effect on the cycle number at fracture. All three of the specimens that included a grid of PAF (Specimens 6, 9, and 12) produced hysteretic response, cyclic load-deformation envelope, strength degradation, and energy dissipation similar to specimens
105
Test 1: RBS24 Test 2: RBS24-PAF12 Test 5: RBS24-PW12
350
0
Energy Dissipated (kN-m-rad)
Specimen
Energy Dissipated (kN-m-rad)
400
Table 5 Summary of moment capacities.
110 Test 10: W36 Test 11: W36-PAF12 Test 12: W36-PAF-array
105 100 95 90 85 80 -5
-4
-3
-2
-1
0
1
2
3
4
Story Drift Ratio (%)
b) W36x150 without RBS Fig. 16. Strength degradation ratios for each group of specimens.
5
[1] Adey BT, Grondin GY, Cheng JJR. Cyclic loading of end plate moment connections. Can J Civ Eng 2000;27(4):683–701. [2] AISC. ANSI/AISC 341-10 seismic provisions for structural steel buildings. Published by the American Institute of Steel Construction; 2010. [3] AISC. ANSI/AISC 358-10 prequalified connections for special and intermediate steel moment frames for seismic applications. Published by the American Institute of Steel Construction; 2010. [4] Beck H, Engelhardt MD. Net section efficiency of steel coupons with power actuated fasteners. ASCE J Struct Eng 2002;128(1):12–21. [5] Chen S-J, Chao YC. Effect of composite action on seismic performance of steel moment connections with reduced beam sections. J Constr Steel Res 2001;57:417–34. [6] Chi B, Uang C-M. Cyclic response and design recommendations of RBS moment connections with deep columns. J Struct Eng ASCE 2002;128(4):464–73. [7] Civjan SA, Engelhardt MD, Gross JL. Retrofit of pre-Northridge moment-resisting connections. J Constr Steel Res ASCE 2000;126(4):445–52. [8] Eatherton MR, Toellner BW, Watkins CE, Abbas E. The effect of powder actuated fasteners on the seismic performance of protected zones in steel moment frames. Virginia Tech SEM Report Series, Report No. CE/VPI-ST-13/05; 2013. [9] Hajjar JF, Leon RT, Gustafson MA, Shield CK. Seismic response of composite momentresisting connections. II: behavior. ASCE J Struct Eng 1998;124(8):877–85. [10] Kosteski N, Packer J, Lecce M. Nailed tubular connections under fatigue loading. J Constr Steel Res 2000;126(11):1258–67.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185 [11] Lee CH, Jeon SW, Kim JH, Uang CH. Effects of panel zone strength and beam web connection method on seismic performance of reduced beam section steel moment connections. J Struct Eng ASCE 2005;131(12):1854–65. [12] Leon RT, Hajjar JF, Gustafson MA. Seismic response of composite moment-resisting connections. I: performance. ASCE J Struct Eng 1998;124(8):868–76. [13] Niessner M, Seeger T. Fatigue strength of structural steel with powder actuated fasteners according to eurocode 3. Stahlbau 1999;68:941–8. [14] Ricles JM, Fisher JW, Lu L-W, Kaufmann EJ. Development of improved welded moment connections for earthquake-resistant design. J Constr Steel Res 2002;58: 565–604. [15] Rogers C, Tremblay R. Inelastic seismic response of frame fasteners for steel roof deck diaphragms. J Constr Steel Res 2003;129(12):1647–57. [16] Toellner B. Evaluating the effect of decking fasteners on the seismic behavior of steel moment frame plastic hinge regions. (M.S. Thesis) Virginia Tech, U.S.; 2013.
185
[17] Tremblay R, Filiatrault A. Seismic performance of steel moment resisting frames retrofitted with a locally reduced beam section connection. Can J Civ Eng 1997;24: 78–89. [18] Uang C-M, Yu Q-S, Noel S, Gross JD. Cyclic testing of steel moment connections rehabilitated with RBS or welded haunch. J Constr Steel Res ASCE 2000;126(1):57–68. [19] Watkins C. Effects of powder actuated fasteners in the protected zone of steel moment frames. Virginia Tech SEM Report Series; 2013. [20] Lee M, Kobayashi G, Tagawa Y. Comparative Study on Effect of Powder Actuated Nail and Stud Welding on Steel Member. Lisbon, Portugal: Proceedings of the 15th World Conference on Earthquake Engineering, September 24-28; 2012. [21] ASTM A6/A6M-12. A6/A6M-12 Standard Specification for General Requirements for Rolled Structural Steel Bars, Plates, Shapes, and Sheet Piling; 2012. [22] ASTM A370-07a. A370-07a Standard Test Methods and Definitions for Mechanical Testing of Steel Products. ASTM; 2007.
Contents lists available at ScienceDirect
Journal of Constructional Steel Research
Experimental investigation on the seismic behavior of steel moment connections with decking attachments B.W. Toellner a, C.E. Watkins b, E.K. Abbas c, M.R. Eatherton c,⁎ a b c
Thornton Tomasetti, 2000 L St. NW, Suite 800, Washington, DC 20036, USA Walter P. Moore, 1301 Mckinney, Suite 1100, Houston, TX 77010, USA Department of Civil and Environmental Engineering, Virginia Tech, 200 Patton Hall, Blacksburg, VA 24061, USA
a r t i c l e
i n f o
Article history: Received 25 June 2014 Accepted 15 November 2014 Available online 15 December 2014 Keywords: Steel moment resisting frames Protected zone Powder actuated fasteners Puddle welds Low cycle fatigue fracture Seismic behavior
a b s t r a c t Steel moment resisting frames rely on large inelastic strains in the beam plastic hinge region to dissipate seismic energy during an earthquake and protect the building against collapse. To limit the potential for premature fracture and because of a lack of test data, fasteners, attachments and defects are prohibited in the plastic hinge region, also referred to as the protected zone in the AISC Seismic Provisions. However, unauthorized attachments and defects occur in many buildings in practice. A set of twelve full-scale moment connection tests were conducted to explore the effect of powder actuated fasteners (PAFs) and puddle welds on the seismic performance of steel moment connections. Both reduced beam section and extended end plate connections were tested with W24 × 62 and W36 × 150 beams. Five specimens included PAFs or puddle welds representing typical steel deck attachment to the top flange of the beam. Three of the specimens included PAFs in a grid over the top and bottom flange and on the web. All twelve specimens passed the qualification criteria for special moment resisting frames (SMRFs) in the AISC Seismic Provisions as they were subjected to a cyclic displacement protocol up to 4% story drift while retaining 80% of their nominal plastic moment capacity. Therefore, the tested moment connection configurations with PAFs and puddle welds were found to produce ductile SMRF type seismic performance. Furthermore, PAFs and puddle welds were found to have negligible effect on cyclic envelope, moment capacity, energy dissipation and strength degradation prior to fracture. © 2014 Elsevier Ltd. All rights reserved.
1. Introduction Special moment resisting frames (SMRFs) depend on large inelastic strains in the beam-to-column connections to dissipate seismic energy and protect buildings from collapse. The ANSI/AISC 341-10 Seismic Provisions for Structural Steel Buildings [2] define the ends of the beams in a SMRF as protected zones at the locations where large inelastic strains are expected. The term “protected” refers to the concept that this portion of the beam should be protected from defects, fasteners, and other discontinuities because their effect on low cycle fatigue fracture is not well understood. For example, AISC 341 [2] prohibits any decking attachments that penetrate the beam flange and any welded, bolted, screwed, or shot-in attachments for other purposes. The extents of the protected zone are defined for each prequalified moment connection included in AISC 358 [3] and are shown graphically in Fig. 1 for two types of SMRF connections. The concept of the protected zone and related restrictions is largely based on a lack of data. In practice, however, the restriction on fasteners in the protected zone creates issues in construction due to the difficulty in communicating and enforcing the requirements. Many instances of ⁎ Corresponding author. E-mail address: [email protected] (M.R. Eatherton).
http://dx.doi.org/10.1016/j.jcsr.2014.11.006 0143-974X/© 2014 Elsevier Ltd. All rights reserved.
unauthorized attachments are submitted to engineers of record each year, and for each reported case of unauthorized attachment, there are likely many more fasteners and defects in the protected zone that go unreported. However, since moment connection tests in the literature focus on the behavior of the connection itself and thus have wellcontrolled defect free plastic hinge regions, there is almost no data on cyclic moment connection behavior with defects or fasteners other than welded shear studs. As a result, engineers of record are faced with a difficult decision between leaving the unauthorized attachment, or requiring an invasive repair method without having any data as to the expected seismic behavior associated with either option. A series of twelve full-scale beam-to-column moment connection tests were conducted to evaluate the effects of powder actuated fasteners (PAFs) and puddle welds applied in the protected zone on the seismic behavior of steel moment connections. The specimens are intended to represent common types of attachments made to the beam in the protected zone such as attachment of steel deck to the beam top flange, attachment of cold-formed steel wall track to the underside of the beam, and attachments for mechanical/electrical nonstructural systems. The objective of the testing program was to investigate whether powder actuated fasteners or puddle welds affected the moment capacity, strength degradation, energy dissipation, and fracture potential of the moment connections as compared to
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
175
To the end of the RBS
Shaded Area is the Protected Zone
RBS Cut a) Reduced Beam Section Moment Connection
b) After Installation
a) Distance equal to the smaller of d or 3bf
Fig. 2. Hilti X-ENP powder actuated fastener.
Flange Width, bf
d Shaded Area is the Protected Zone
b) Unstiffened End Plate Moment Connection Fig. 1. Examples of protected zone for two moment connection types.
specimens with no fasteners. The special moment resisting frame (SMRF) qualification criteria of AISC 341-10 [2] were used as a benchmark for defining acceptance as the specimens were considered to produce sufficiently ductile behavior if they could retain a significant portion of their moment capacity after being subjected to cyclic loading up to 4% story drift. 2. Background There have been many experimental programs on full-scale moment connections since the 1994 Northridge Earthquake and the 1995 Kobe Earthquake exposed unexpected fractures in steel beam to column connections. Experimental programs focused on a range of topics including reduced beam section connections (e.g. [6]), extended end plate connections (e.g. [1]), effect of composite slabs (e.g. [9]), panel zone strength (e.g. [11]), and much more. A subset of these testing programs included a composite slab with welded shear studs attached to the beam in the protected zone (e.g. [5,7,9,12,14,18]). Out of this body of tests, one beam-to-column connection specimen with welded shear studs in the protected zone experienced fracture of the flange initiating at the shear stud [14]. It was concluded that the fracture of the beam flange was a direct result of the reduced notch toughness of the base metal caused by shear stud welding. The other specimens in this testing program and specimens from other testing programs did not develop fracture at the welded stud location. Besides the effect of the fastener itself on stress concentrations and material properties, connections between the concrete slab and the steel beam strongly affect the strain distribution in the beam. Beamto-column tests using composite slabs and welded shear studs through the protected zone have demonstrated that composite action in the connection region causes significantly increased strain demands in the beam bottom flange [9,12,17]. Both the rapid change from composite section to bare steel beam right at the connection to the column and larger strain demands in the critical bottom flange where the majority of fractures were observed after the Northridge earthquake were considered detrimental.
However, the effect of a particular type of fastener on stress concentrations, base metal properties, and degree of composite action is heavily dependent on the type and configuration of fasteners. For this reason, it is necessary to investigate the effect of each type of artifact (i.e. fastener or defect) on moment connection seismic behavior separately. Powder actuated fasteners (PAFs) are a common method for attaching steel deck to the top flange of steel beams, as well as a common means for attaching nonstructural elements such as cold-formed steel tracks and mechanical, electrical, and plumbing elements to the structure. The fastener (example shown in Fig. 2) is driven into the steel plies using a fastening tool. The effect of PAFs on fracture of steel coupons subjected to monotonic tension has been studied by Beck and Engelhardt [4] and it was determined that coupons with PAFs had higher tensile strength than coupons with equivalent drilled holes. Moreover, a typical coupon specimen with PAF was shown to reach a strain of 17% before fracture which is less than a specimen with no holes, but approximately 70% more ductility than a similar specimen with drilled holes. The increased strength and ductility of steel coupons with PAFs as compared to drilled holes might be attributed to increased strength in the surrounding base metal caused during the fastener application or to residual compressive stresses in the material surrounding the hole. Studies have also been conducted to investigate the high cycle fatigue performance of steel with PAFs (e.g. [13]), tube connections using PAF (e.g. [10]), and steel deck attachment using PAF (e.g. [15]). The extent of material affected by PAF was investigated [20] in which it was found that a particular type of PAF affected the Vickers hardness up to 6 mm from the edge of the fastener. However, the behavior of steel with PAFs subjected to cycles of large inelastic strains, such as those experienced in moment frame plastic hinges, has not been experimentally investigated. 3. Testing program description 3.1. Test setup The test setup is shown in Figs. 3 and 4. The configuration uses a W14 × 257 vertical column, approximately 3.67 m tall that is restrained against lateral translation at the top and bottom. The W24 × 62 and W36 × 150 beam specimens attach to the column with a bolted end plate connection to allow the same column to be used for every test. Furthermore, the two ends of the beam were used for two different tests to maximize the efficiency of steel used for these tests. Additional details on geometry and reaction frames are given in Toellner [16] and Watkins [19]. Load was applied with an MTS model 201.70 actuator with force capacity of 956 kN in tension and 1468 kN in compression and total stroke equal to 508 mm. The geometry of the test setup is intended to simulate an exterior column subassemblage of a moment frame undergoing rotation due to
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
W14x257 Column
3670 mm
176
MTS servo-controlled actuator: 965 kN capacity and 508 mm stroke P= Actuator Force
Lateral bracing locations
Beam CL
Table 1 Cyclic displacement protocol. Percent story drift (%)
Number of cycles
Displacement at actuator (mm)
0.38 0.50 0.75 1.00 1.50 2.00 3.00 4.00 4.66
6 6 6 4 2 2 2 2 Up to 5
19 26 39 52 78 104 155 207 241
1543mm for W24x62 1359 mm for W36x150
Beam rotated and flipped for subsequent specimen Lf = 4973 mm L c = 5182 mm
Fig. 3. Test frame layout.
story drift. The 5.18 m distance from the column centerline to the actuator represents the distance from the column centerline to the beam inflection point in a 10.36 m long bay assuming an idealized moment diagram. Similarly, the column height of 3.67 m represents the distance between column inflection points on adjacent floors assuming an idealized moment distribution and implying a building with 3.67 m floor heights. A third reaction frame provided lateral restraint to the test specimens. The location of the frame could be moved depending on unbraced length requirements of the specimens. The resulting lateral bracing points were 0.84 m, 1.49 m, and 4.65 m from the face off the column for W24 × 62 beam specimens and 1.14 m, 1.79 m, and 4.65 m for the W36 × 150 beam specimens. 3.2. Loading protocol and qualification criteria The specimens were subjected to the cyclic story drift history described in AISC 341-10 for special moment frames [2] and given in Table 1. The target displacements at the actuator location were determined by multiplying the target story drifts by the distance between the actuator and the column centerline, Lc = 5.18 m. Beyond the qualification amplitude of 4% story drift, additional cycles were performed at an amplitude of 4.7% story drift which corresponds to the maximum stroke of the actuator. The tests were displacement controlled using
Fig. 4. Photograph of test setup.
external displacement feedback from a string potentiometer located below the actuator with a displacement rate of 61 mm/min for all tests. The SMRF qualification criteria from AISC 341-10 [2] were used as a benchmark to determine whether the connections containing PAFs or puddle welds achieved sufficient ductility to produce SMRF type of performance with the desirable seismic behavior implied therein. The SMRF qualification criteria in AISC 341-10 chapter K specify that a specimen must undergo one complete cycle at 4% story drift while sustaining a moment resistance at the face of the column that is at least 80% of the nominal plastic moment capacity. The nominal plastic moment capacity is computed with nominal yield stress and nominal plastic section modulus.
3.3. Test specimens The test program consisted of twelve specimens as listed in Table 2. This set of specimens includes variation in beam depth (nominal W24 to W36), flange thickness (15 mm to 24 mm), flange width (178 mm to 305 mm), and beam weight (92 kg/m and 223 kg/m). The flange thickness may be an important variable as the location of the PAF tip in the depth of the flange may affect the potential for fracture initiation and both the flange thickness and width will affect the relative magnitude of the disturbance in the strain field. The depth of the section will affect the inelastic strain demands as the distance from the neutral axis to the fastener will vary for the same curvature demand. The flange width and incorporation of RBS cuts will further affect inelastic strain demands at the flange tips and thus are also key variables that were varied in this study. Six of the specimens included a reduced beam section (RBS), but also utilized an extended end plate so that the same column could be reused. It is noted that while RBS specimens don't typically include endplates, it is expected that the relatively rigid endplate did not have a significant effect on the inelastic deformations of the RBS plastic hinge which are the focus of this investigation. The end plate connections were designed in accordance with ANSI/AISC 358-10 for the full strength of the beam without RBS and additional details can be found in Eatherton et al. [8]. A similar variation of fasteners was used for both beam depths as shown in Fig. 5. Fasteners at 305 mm spacing are typical for deck attachments to the top flange or partition attachments to the underside of the bottom flange so both PAF and puddle weld configurations with this spacing were included in the testing program. Three specimens (Specimens 6, 9, and 12) included a grid of PAFs on both flanges with spacing equal to 51 mm, equally spaced lines of PAF across the width of the flange with spacing not less than 51 mm, and distance from center of PAF to the edge of the flange of 25 mm. Both Specimen 9 and Specimen 12 also incorporated a grid of PAFs on one side of the web as depicted in Fig. 5m. Although the specimens with a grid of PAF don't represent attachments typically made in practice, the grid specimens were intended to identify the worst case locations as PAF could conceivably be located in any of these positions. Assuming negligible interaction between the fasteners, a single PAF grid test can therefore represent a multitude of individual tests in which a PAF is present at any of the given locations.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
177
Table 2 Matrix of test specimens. Specimen
Beam size
1 2 5b 3b 4 6 7 8 9 10 11 12
W24 × 62 W24 × 62 W24 × 62 W24 × 62 W24 × 62 W24 × 62 W36 × 150 W36 × 150 W36 × 150 W36 × 150 W36 × 150 W36 × 150
Connection type
Protected zone lengtha (mm)
Nominal Mp and 0.8Mp (kn m)
RBS RBS RBS BUEEP BUEEP BUEEP RBS RBS RBS BSEEP BSEEP BSEEP
620 620 572 572 620 572 1003 1003 1003 818 818 818
556/445 556/445 556/445 865/691 865/691 865/691 2261/1810 2261/1810 2261/1810 3281/2626 3281/2626 3281/2626
Fasteners in the protected zone None 4 PAFs spaced at 305 mm 4 puddle welds spaced at 305 mm None 4 PAFs spaced at 305 mm PAF grid None 4 puddle welds spaced at 305 mm PAF grid None 4 PAFs spaced at 305 mm PAF grid
RBS = reduced beam section. BUEEP = bolted unstiffened extended end plate. BSEEP = bolted stiffened extended end plate. a Protected zone length measured from column face. b Specimen 3 conducted prior to Specimen 5, but reordered for test matrix.
fastener size was chosen to represent the largest size PAF used in deck attachment to steel beams. The arc spot puddle welds were applied by a qualified lab technician who is certified for this type of welding and other types of welding by the American Welding Society. The puddle welds were 19 mm in diameter and made through a small square of 20 gage metal decking and then the excess decking material was trimmed off using a torch to allow strain gage application.
It is noted however, that the grid of PAF on one specimen is an extreme condition that is not indicative of construction practice. 3.4. Power actuated fasteners and puddle weld information The PAFs used in this study were Hilti X-ENP-19 L15 nails driven into test specimens with a 0.27 caliber Hilti DX 76 powder-actuated tool. The
762
127
458
203
2 Spaces at 304
247
337
No Fasteners No Fasteners
a) Specimen 1 54
3 Spaces at 304
2 Spaces at 304 Powder Actuated Fastener (PAF)
PAF
g) Specimen 7 247
k) Specimen 11 337
229
51 Typical
b) Specimen 2
No Fasteners Puddle Welds
c) Specimen 3
l) Specimen 12
54
3 Spaces at 304
51 Typical
PAF
889
h) Specimen 8
229
127
PAF
d) Specimen 4 54
3 Spaces at 304
PAF
991
Puddle Welds
PAF
152 Typical
i) Specimen 9
e) Specimen 5
889
51
51 Typical
No Fasteners
m) PAF Applied to One Side of
711
f) Specimen 6
j) Specimen 10 PAF Fig. 5. Specimen configurations.
Web for Specimens 9 and 12
178
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
4. Test results
Table 3 Measured steel properties from tension coupons.
W24 × 62
W36 × 150
Yield stress, Fy (MPa)
Ultimate stress, Fu (MPa)
Elongation at fracture using punch marks (%)
363 365 360 363 364 374 376 372
489 478 461 476 479 470 478 476
22 27 23 24 30 30 29 30
Top flange Bottom flange Web Average Top flange Bottom flange Web Average
3.5. Ancillary material tests All beam specimens were specified to be ASTM A992 structural steel with a minimum yield stress of 345 MPa and a minimum tensile strength of 448 MPa. All beams of the same size were obtained from the same heat of material. Tension tests were performed on six coupon specimens cut with longitudinal orientation from the top flange, bottom flange, and web from the undeformed regions of one W24 × 62 beam and one W36 × 150 beam in accordance with ASTM A6/A6M-12 [21]. The geometry of the coupon specimens was in accordance with ASTM 370-07a [22] with a width of 38 mm, a total length of 457 mm and a gage length of 203 mm. Each specimen included punch marks at the ends of the gage length for which the distance between punch marks was measured before and after testing. The yield stress, ultimate stress, and elongation at fracture as measured using the distance between punch marks are given in Table 3.
All twelve specimens tested in this experimental program satisfied the SMRF qualification criteria of AISC 341-10. This means that moment connections with configurations of fasteners such as shown in Fig. 5 can produce SMRF type of behavior with all the ductility and energy dissipation implied therein. The results of each group of tests are presented in this section, and then the results are compared and synthesized in the following section. The plots of load-deformation presented in this section show the moment at the face of the column given by the applied force, P, multiplied by the distance to the face of the column, Lf (see Fig. 3) and normalized by the nominal plastic moment capacity, Mp, calculated using nominal section properties and nominal yield stress. The horizontal axes are story drift calculated as the displacement at the actuator, δSP2 (see Fig. 6) divided by the distance to the column centerline, Lc (see Fig. 3).
4.1. Behavior of W24 × 62 specimens with RBS The progression of limit states was similar for all specimens in this group starting with significant yielding in the extreme fibers at the reduced section, spread of plasticity, local buckling of the flanges in association with out-of-plane buckling of the web, crack initiation typically at the flange tips on the inside face of the local buckle, and fracture propagation through an entire flange. All specimens experienced some amount of fracturing at the minimum flange section. In all cases, the fractures initiated at the edge of the flange on the inside of a local buckle and propagated toward the center of the beam flange. Fig. 7 shows the
Load Cell and Displacement Sensor in the Actuator
LVDT 8 Inclinometer 1 Inclinometer 2 Inclinometer 3
LVDT 1 LVDT 5 LVDT 6 LVDT 2
LVDT 4
Story Drift, δ SP2 / Lc (%)
Flange Fracture
LVDT 9 LVDT 7
String Potentiometer 3
LVDT 3
Flange Fracture
a) Specimen 1, No Fasteners
Applied Moment , PLf / Mp
The instrumentation plan was designed to include enough displacement sensors to decompose the applied story drift into components due to the reaction frame, column outside the panel zone, column panel zone, end plate, plastic hinge, and elastic deformations of the beam outside the plastic hinge region. Nine LVDTs, three string potentiometers, and three inclinometers were used as shown in Fig. 6. Additional details are included in Eatherton et al. [8]. In addition to displacement and rotation sensors, a set of high strain rated strain gages were applied in configurations across the flange width with the intent to investigate differences in the strain gradient with and without fasteners. The strain gages had a published strain range of 10% to 15% and were applied in a line of seven strain gages on the outside of the flange and four on the inside face of the flange. In locations with a fastener at the center of the flange, only six strain gages were applied on the outside of the flange.
Applied Moment, PLf / Mp
3.6. Instrumentation
String Potentiometer 2 δSP2
Note: LVDT = Linear Variable Differential Transformer
Fig. 6. Instrumentation plan.
Story Drift , δ SP2 / Lc (%)
b) Specimen 2, PAF at 305 mm Fig. 7. Load-deformation behavior of W24 × 62 specimens with RBS.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
load-deformation behavior for Specimen 1 without any fasteners, and Specimen 2 with PAF at 305 mm spacing. Both specimens experienced fracture during the second cycle at 4.7% story drift and Specimen 5 (not shown) experienced fracture in the third cycle at 4.7% story drift. The strength degradation was slightly delayed for Specimen 2, but this was likely due to slight differences in the specimen geometry and material. The load-deformation response for Specimen 5 was similar and thus not shown. The fracture surfaces were analyzed to determine the crack initiation locations and fracture propagation paths. Fig. 8c and d shows that the fracture in Specimen 1 with no fasteners started at the flange tips, with initiation occurring at the inside of the local buckles. Crack initiation at the flange tips was also noted in Specimen 2 in which the fracture propagated toward the center of the section displaying five distinct cycles of ductile fracture propagation before the crack reached the PAF location and brittle overload occurred (see Fig. 8a). Although a similar process occurred in Specimen 5 including crack initiation, limited ductile fracture propagation, and brittle overload as shown in Fig. 8b, there was less ductile fracture propagation than Specimen 2. It is noted that the fracture in Specimen 5 occurred in the unadulterated flange opposite from the one that included puddle welds. 4.2. Behavior of W24 × 62 specimens without RBS The progression of limit states was similar for all specimens in this group starting with significant yielding in the extreme fibers near the endplate connection, spread of plasticity, and local buckling of the flanges in association with out-of-plane buckling of the web. None of the specimens experienced significant fractures even after being subjected to multiple cycles at 4.7% story drift.
Ductile Overload
Cycle Marks
PAF Location
179
Fig. 9 shows relatively similar load-deformation behavior between Specimen 3 which had no fasteners and Specimen 6 which had a grid of PAF on both flanges as shown in Fig. 5f. Specimen 4 with PAF at 305 mm exhibited behavior similar to Specimen 3 and is thus not discussed further here. See Eatherton et al. [8] for more information. Testing of Specimen 3 (as well as Specimen 4 which is not shown) was arbitrarily stopped after 2 cycles at 4.7% story drift, but no fractures were observed. Testing of Specimen 6 was continued for five cycles at 4.7% story drift until such time as the load-deformation behavior appeared to be stabilizing, implying that the inelastic local buckling (see Fig. 10) was not varying much from one cycle to the next. At a few of the PAF locations at the peaks of the local buckles, the hole elongated and small tears formed radiating out from the PAF (see Fig. 10). As compared to the cracks that initiated at the flange tips of the RBS in the previous group of specimens, these tears appeared to propagate very slowly, extending approximately 10 mm from the edge of the PAF by the end of the test. 4.3. Behavior of W36 × 150 specimens with RBS The W36 × 150 specimens experienced a similar progression of limit states as the W24 × 62 specimens starting with yielding of the extreme fibers at the reduced section, spread of plasticity, local buckling of the flange in association with out-of-plane buckling of the web, crack initiation, and fracture propagation. However, unlike the W24 × 62 specimens, the cracks initiated on the face of the flanges at the inside of the local buckles as shown in Fig. 11a. For Specimen 7, the cracks were identified during the 4% story drift cycles and then during the 4.7% story drift cycles, fracture propagated through the flange and web as shown in Fig. 11c. A similar process occurred for Specimen 9 that had a grid of
Critical Flow Region 2 1
Crack Initiation and Slow Growth
Final Overload Region
Crack Initiation
Shear Lips
Brittle Overload
a) Specimen 2 Failure Origin
c) Specimen 1
Crack Initiation Shear Lips
b) Specimen 5 Internal Cracks
d) Specimen 1
Fig. 8. Analysis of fracture surfaces for W24 × 62 specimens with RBS.
Failure Origin
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
Applied Moment , PL f / Mp
180
cycle at 4% story drift. Specimen 8, which is not shown here, experienced a fracture at the toe of the stiffener which was unrelated to the puddle welds that were present on that specimen. As shown in Fig. 12, the load-deformation behavior of the specimen without fasteners, and the specimen with a grid of PAF on both flanges was quite similar prior to fracture. The differences in the peak strength and strength degradation will be analyzed further in a subsequent section.
Test Stopped Without Fracture
4.4. Behavior of W36 × 150 specimens without RBS
Story Drift , δ SP2 / Lc (%)
Applied Moment , PL f / M p
a) Specimen 3, No Fasteners Test Stopped Without Fracture
The progression of limit states for the W36 × 150 specimens that didn't include an RBS was the same as those for the W24 × 62 specimens with no RBS. Although some tearing was observed at the stiffener toe, Specimens 10 and 11 with no fasteners and PAF at 305 mm spacing respectively, did not experience significant fracture as they were subjected to the full loading protocol including five cycles at 4.7% story drift. Specimen 12 that included a grid of PAF over both flanges and web, experienced a fracture during the fourth cycle at 4.7% story drift. However, the load-deformation behavior of Specimens 10 and 12 was quite similar as shown in Fig. 13. Specimen 12 underwent substantial local buckling and inelastic strains prior to fracture as shown in Fig. 14a. The fracture in specimen 12, shown in Fig. 14b, initiated on the surface of the flange at a PAF located on the inside of a local buckle. 5. Discussion of results
Story Drift, δ SP2 / L c (%)
b) Specimen 6, Grid of PAF
This section summarizes several studies performed to compare test data in order to identify differences between the performance of control specimens and those with PAFs or puddle welds. These studies include comparisons of hysteretic behavior, strength degradation and energy dissipation.
Fig. 9. Load-deformation behavior of W24 × 62 specimens without RBS.
5.1. Fracture behavior PAF over both flanges. As shown in Fig. 11b, the fracture occurred at the reduced section where the local buckle had the largest curvature. An analysis of the fracture surface, shown in Fig. 11d, revealed that the crack initiated at the surface of the flange around a PAF at the inside of the local buckle. These cracks were observed to form during the second
Tear
Fig. 10. Specimen 6 at the end of the test.
Several general observations are made regarding the fracture behavior of the twelve specimens. Table 4 provides a summary of the fractures and shows that all six RBS specimens experienced some degree of fracture during 4.7% story drift cycling. Specimen 1 experienced a ductile flange tear and the other five RBS specimens experienced brittle fractures along an entire flange width and a portion of the web. The only non-RBS specimen to experience a fracture was Specimen 12 which contained a grid of PAF over both flanges and the web. Strain gage data (not shown here, see Eatherton et al. [8]) did not reveal noticeable differences in flange strain gradient between specimens with PAF, puddle welds or no fasteners. At large story drifts, inelastic strains were found to be dominated by curvature related to local buckles and inelastic strain was not found to concentrate in the vicinity of PAF or puddle welds. For the W24 × 62 beam specimens with RBS, the fractures occurred during the 2nd or 3rd cycle at 4.7% story drift regardless of whether there were PAF, puddle welds, or no fasteners included in the protected zone. This suggests that although the PAF may interact with the fracture process, the configurations of PAF and puddle welds tested did not cause fracture significantly earlier than specimens with no fasteners. The W36 × 150 RBS specimen with no fasteners fractured during the 5th cycle at 4.7% story drift, whereas the specimens with a grid of PAF and puddle welds fractured in the 2nd and 1st cycles at 4.7% story drift respectively. However the fracture in the specimen with puddle welds (Specimen 8) was not related to the puddle welds, but instead occurred at the junction of the stiffener and flange. This implies some inherent variability in the cycle at which fracture occurs. Specimen 12, which included a grid of PAF, fractured during the 4th cycle at 4.7% story drift whereas the other non-RBS specimens (Specimens 10 and
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
a) Specimen 7
181
b) Specimen 9 Crack Initiation
Brittle Overload Propagation Direction
Critical Flow Region Shear Lips
c) Specimen 7 Fracture Surface Crack Initiation
Powder Actuated Fastener
Brittle Overload Ductile Overload and Critical Flow Region
Shear Lips
d) Specimen 9 Fracture Surface Fig. 11. Fracture for Specimen 7 without fasteners and Specimen 9 with grid of PAF.
11) did not experience fracture through five cycles at 4.7% story drift even though Specimen 11 included PAF at 305 mm spacing. The results for the W36 × 150 specimens that included a grid of fasteners suggest that one or more PAF applied in critical locations such as the inside of a local buckle may lead to fracture during an earlier cycle than specimens without fasteners. The potential for fracture is supported by the small tears that formed at PAF in critical locations after a large number of inelastic strain cycles. It is noted, however, that fracture occurred after SMRF qualification had been satisfied, the difference in the cycle number at fracture was small, and that the two W36 × 150 PAF specimens that fractured included a grid of PAF that represents an extreme condition, not indicative of construction practice. Since the sample size was small and the difference in the cycle number at fracture
was small, it was not possible to evaluate the difference in a statistically significant manner.
5.2. Load-deformation behavior Moment-drift envelopes for all tests are shown in Fig. 15. The envelope was developed by extracting the peak moments for any cycle at each story drift level. It is shown in Fig. 15a that the two groups of W24 × 62 specimens had nearly identical envelopes with slight variations that were likely not related to the puddle welds or PAF. Similarly, Fig. 15b shows that the cyclic envelopes for the W36 × 150 specimens were not affected by inclusion of puddle welds or PAF. It is expected
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
Applied Moment, PL f / M p
Flange Fracture
Applied Moment, PLf / Mp
182
Test Stopped Without Fracture
Story Drift , δ SP2 / Lc (%)
Story Drift , δ SP2 / Lc (%)
Flange Fracture
a) Specimen 10 without Fasteners
Applied Moment, PLf / Mp
Applied Moment , PL f / Mp
a) Specimen 7 without Fasteners
Flange Fracture
Story Drift , δ SP2 / Lc (%)
b) Specimen 9 with Grid of PAF Fig. 12. Load-deformation behavior of W36 × 150 specimens with RBS.
that the accelerated strength degradation of Specimen 11 was due to larger initial imperfections, not the PAF. The similarity between the envelopes with and without fasteners in the protected zone is also demonstrated in Table 5. The maximum moment capacities are all within 1% and 5% of the average value for their group for W24 × 62 and W36 × 150 specimens, respectively. Similarly, the moment capacities at the qualification cycle were within 5% of their group average for all groups except the W36 × 150 specimens with no RBS which experienced more variability with as much as 12% difference from the average moment capacity. Table 5 also shows that all twelve specimens satisfied the AISC 34110 qualification criteria for special moment resisting frames by retaining 80% of the nominal moment capacity, Mp, through one full cycle at 4% story drift. Furthermore, the moment capacity during the 4% story drift qualification cycle was on average 37% greater than the 0.8Mp.
Story Drift , δ SP2 / Lc (%)
b) Specimen 12 with Grid of PAF Fig. 13. Load-deformation behavior of W36 × 150 specimens without RBS.
the non-RBS specimens showed strength degradation ratios that increased as compared to the 4% cycles implying that local buckling deformations began to stabilize. 5.4. Energy dissipation The energy dissipation behavior of each specimen was also compared for each group of specimens. Energy dissipation was calculated using the trapezoidal rule for numerical integration and the resulting values for specimens with RBS are shown in Fig. 17. At the conclusion of the 4% story drift cycles, the energy dissipation for the specimens with PAF or puddle welds was on average within 3% of the energy dissipated by the associated specimens without fasteners. The energy dissipation at the end of the test varied more as it was highly dependent on the number of cycles performed at 4.7% story drift level.
5.3. Strength degradation 6. Summary and conclusions A strength degradation ratio, η, was defined for a given story drift amplitude, i, as the ratio of the moment at the second peak displacement to the moment at the first peak displacement. Despite its designation as a strength degradation ratio, values greater than 100% are possible and indicate an increase in moment capacity due to strain hardening. Fig. 16 shows the strength degradation ratio for the specimens without RBS. It is shown that within each group, the specimens with and without fasteners exhibit very similar strength degradation trends. For RBS specimens (not shown here), strain hardening controls through the 2% story drift cycles, followed by strength degradation associated with local buckling. Specimens without RBS showed strain hardening through the 3% story drift cycles suggesting that local buckling was delayed as compared to RBS specimens. During the 4.7% story drift cycles,
Twelve full-scale beam-to-column moment connection tests were conducted with a range of variables to investigate the effect of powder actuated fasteners and puddle welds on the seismic behavior of moment frames. Variations in the test specimens included control specimens with no fasteners, PAF at 305 mm spacing representing typical deck attachment, puddle welds at 305 mm spacing, and PAF in a dense grid over the plastic hinge region with 25 mm spacing to any edge. Sections tested included W24 × 62 and W36 × 150 beams with flange thicknesses of 15 mm and 24 mm, respectively. Reduced beam section (RBS) connections and non-RBS connections were considered representative of the range of currently prequalified special moment resisting frame connection types.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
183
1000 Test Test Test Test Test Test
800
Applied Moment (kN-m)
600 400
1: 2: 5: 3: 4: 6:
RBS24 RBS24-PAF12 RBS24-PW12 W24 W24-PAF12 W24-PAF-array
200 0 -200 -400 -600 -800 -1000 -5
-4
-3
-2
-1
0
1
2
3
4
5
3
4
5
Story Drift (%)
a) W24x62 Envelopes 5000 Test Test Test Test Test Test
4000
Crack Initiation
Shear Lips
3000
Applied Moment (kN-m)
a) Specimen 12 Buckled Shape
2000
7: RBS36 8: RBS36-PW12 9: RBS36-PAF-array 10: W36 11: W36-PAF12 12: W36-PAF-array
1000 0 -1000 -2000 -3000
Brittle Overload
-4000 -5000 -5
-4
-3
-2
-1
0
1
2
Story Drift (%)
Powder Actuated Fasteners
b) W36x150 Envelopes Fig. 15. Moment vs. story drift envelopes extracted from cyclic test data.
b) Specimen 12 Fracture Surface Fig. 14. Behavior of Specimen 12 with grid of PAF over both flanges and the web.
All twelve specimens satisfied the SMRF qualification criteria in ANSI/AISC 341-10 Chapter K, possessed significant ductility, demonstrated substantial energy dissipation, and according to ANSI/AISC 3410-10 would be considered adequate for use in a special moment resisting frame. The behavior of the twelve specimens was quantitatively compared using several different approaches. The cyclic-load deformation behavior and cyclic envelopes were compared for specimens with no fasteners, puddle welds, and PAF. It was found that the fasteners
had negligible effect on the hysteretic behavior or envelope prior to fracture. Furthermore the achieved moment capacities during the qualification cycle and maximum moment capacities were unchanged with the inclusion of puddle welds and PAF. The effect of puddle welds and PAF on strength degradation and energy dissipation was similarly found to be negligible prior to fracture, even in the specimens with a dense grid of PAF. Although the W24 × 62 specimens suggest that the PAF and puddle welds did not change which cycle saw fracture, the results for the two W36 × 150 specimens that included a grid of fasteners suggest that
Table 4 Summary of fracture cycle and location. Specimen number
Specimen description
Cycle at significant fracture
Fracture location
1 2 5 3 4 6 7 8 9 10 11 12
W24 W24 W24 W24 W24 W24 W36 W36 W36 W36 W36 W36
2nd cycle at 4.7% story drift 2nd cycle at 4.7% story drift 3rd cycle at 4.7% story drift N/A N/A N/A 5th cycle at 4.7% story drift 2nd cycle at 4.7% story drift 1st cycle at 4.7% story drift N/A N/A During 4th cycle at 4.7% story drift
Bottom flange, near center of RBS Through PAF location at center of RBS on top flange Bottom flange, away from puddle welds, near center of RBS Test stopped after two cycles at 4.7% No fracture Test stopped after two cycles at 4.7% No fracture Test stopped after five cycles at 4.7% No fracture Top flange at center of RBS, initial fracture partial depth then opened Bottom flange at tip of end-plate stiffener, not at puddle welds Top flange near center of RBS Through two PAFs Test stopped after five cycles at 4.7% No fracture Test stopped after five cycles at 4.7% No fracture Bottom flange through multiple PAF and up to a PAF located on the web
× × × × × × × × × × × ×
62 with RBS but no fasteners 62 with RBS and PAF at 305 62 with RBS and puddle welds at 305 62 without RBS or fasteners 62 without RBS and PAF at 305 62 without RBS and grid of PAF 150 with RBS and no fasteners 150 with RBS and puddle welds at 305 150 with RBS and grid of PAF 150 without RBS and no fasteners 150 without RBS and PAF at 305 150 without RBS and grid of PAF
184
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185
W24 × 62
RBS
1 2
No RBS
W36 × 150
RBS
No RBS
Fasteners
None PAF at 305 mm Puddle welds None PAF at 305 mm Grid of PAF None Puddle welds Grid of PAF None PAF at 305 mm Grid of PAF
5 3 4 6 7 8 9 10 11 12
Max. moment capacity (kN m)
Moment capacity 4% drift (kN m)
710 705
545 586
701 1007 994
548 923 933
998 3196 2943 3045 4131 3856
961 2835 2756 2698 3734 3091
4108
3638
Strength Degredation Ratio (%)
250 200 150 100 50 0
1
2
3
4
5
a) W24x62 with RBS 2500 Test 7: RBS36 Test 8: RBS36-PW12 Test 9: RBS36-PAF-array
2000
1500
1000
500
0
0
1
2
3
4
5
Story Drift (%)
b) W36x150 with RBS Fig. 17. Energy dissipation for each group of specimens.
with no fasteners through the SMRF qualification cycles. Since all twelve specimens satisfied the qualification criteria, it has been demonstrated that moment connections with puddle welds and PAF in configurations such as those tested in this study are capable of substantial ductility consistent with SMRF behavior.
110 Test 3: W24 Test 4: W24-PAF12 Test 6: W24-PAF-array
100 95
Acknowledgments
90
This material is based on work supported by the Hilti, Inc. and the American Institute of Steel Construction. In-kind funding was provided by the Banker Steel and Applied Bolting Technology.
85 80 -5
-4
-3
-2
-1
0
1
2
3
4
5
Story Drift Ratio (%)
References
a) W24x62 without RBS Strength Degredation Ratio (%)
300
Story Drift (%)
one or more PAF applied in critical locations (i.e. the inside of a local buckle) may lead to fracture during an earlier cycle in the displacement protocol. In the three specimens with grids of fasteners (Specimens 6, 9 and 12), tears initiated at PAF in critical locations. It is noted, however, that fracture occurred after SMRF qualification had been satisfied, the difference in the cycle number at fracture was small, and that the two W36 × 150 specimens that fractured included a grid of PAF that represents an extreme condition, not typical of construction practice. On the other hand, PAF or puddle welds applied in a pattern typical of deck attachment appeared to have no effect on the cycle number at fracture. All three of the specimens that included a grid of PAF (Specimens 6, 9, and 12) produced hysteretic response, cyclic load-deformation envelope, strength degradation, and energy dissipation similar to specimens
105
Test 1: RBS24 Test 2: RBS24-PAF12 Test 5: RBS24-PW12
350
0
Energy Dissipated (kN-m-rad)
Specimen
Energy Dissipated (kN-m-rad)
400
Table 5 Summary of moment capacities.
110 Test 10: W36 Test 11: W36-PAF12 Test 12: W36-PAF-array
105 100 95 90 85 80 -5
-4
-3
-2
-1
0
1
2
3
4
Story Drift Ratio (%)
b) W36x150 without RBS Fig. 16. Strength degradation ratios for each group of specimens.
5
[1] Adey BT, Grondin GY, Cheng JJR. Cyclic loading of end plate moment connections. Can J Civ Eng 2000;27(4):683–701. [2] AISC. ANSI/AISC 341-10 seismic provisions for structural steel buildings. Published by the American Institute of Steel Construction; 2010. [3] AISC. ANSI/AISC 358-10 prequalified connections for special and intermediate steel moment frames for seismic applications. Published by the American Institute of Steel Construction; 2010. [4] Beck H, Engelhardt MD. Net section efficiency of steel coupons with power actuated fasteners. ASCE J Struct Eng 2002;128(1):12–21. [5] Chen S-J, Chao YC. Effect of composite action on seismic performance of steel moment connections with reduced beam sections. J Constr Steel Res 2001;57:417–34. [6] Chi B, Uang C-M. Cyclic response and design recommendations of RBS moment connections with deep columns. J Struct Eng ASCE 2002;128(4):464–73. [7] Civjan SA, Engelhardt MD, Gross JL. Retrofit of pre-Northridge moment-resisting connections. J Constr Steel Res ASCE 2000;126(4):445–52. [8] Eatherton MR, Toellner BW, Watkins CE, Abbas E. The effect of powder actuated fasteners on the seismic performance of protected zones in steel moment frames. Virginia Tech SEM Report Series, Report No. CE/VPI-ST-13/05; 2013. [9] Hajjar JF, Leon RT, Gustafson MA, Shield CK. Seismic response of composite momentresisting connections. II: behavior. ASCE J Struct Eng 1998;124(8):877–85. [10] Kosteski N, Packer J, Lecce M. Nailed tubular connections under fatigue loading. J Constr Steel Res 2000;126(11):1258–67.
B.W. Toellner et al. / Journal of Constructional Steel Research 105 (2015) 174–185 [11] Lee CH, Jeon SW, Kim JH, Uang CH. Effects of panel zone strength and beam web connection method on seismic performance of reduced beam section steel moment connections. J Struct Eng ASCE 2005;131(12):1854–65. [12] Leon RT, Hajjar JF, Gustafson MA. Seismic response of composite moment-resisting connections. I: performance. ASCE J Struct Eng 1998;124(8):868–76. [13] Niessner M, Seeger T. Fatigue strength of structural steel with powder actuated fasteners according to eurocode 3. Stahlbau 1999;68:941–8. [14] Ricles JM, Fisher JW, Lu L-W, Kaufmann EJ. Development of improved welded moment connections for earthquake-resistant design. J Constr Steel Res 2002;58: 565–604. [15] Rogers C, Tremblay R. Inelastic seismic response of frame fasteners for steel roof deck diaphragms. J Constr Steel Res 2003;129(12):1647–57. [16] Toellner B. Evaluating the effect of decking fasteners on the seismic behavior of steel moment frame plastic hinge regions. (M.S. Thesis) Virginia Tech, U.S.; 2013.
185
[17] Tremblay R, Filiatrault A. Seismic performance of steel moment resisting frames retrofitted with a locally reduced beam section connection. Can J Civ Eng 1997;24: 78–89. [18] Uang C-M, Yu Q-S, Noel S, Gross JD. Cyclic testing of steel moment connections rehabilitated with RBS or welded haunch. J Constr Steel Res ASCE 2000;126(1):57–68. [19] Watkins C. Effects of powder actuated fasteners in the protected zone of steel moment frames. Virginia Tech SEM Report Series; 2013. [20] Lee M, Kobayashi G, Tagawa Y. Comparative Study on Effect of Powder Actuated Nail and Stud Welding on Steel Member. Lisbon, Portugal: Proceedings of the 15th World Conference on Earthquake Engineering, September 24-28; 2012. [21] ASTM A6/A6M-12. A6/A6M-12 Standard Specification for General Requirements for Rolled Structural Steel Bars, Plates, Shapes, and Sheet Piling; 2012. [22] ASTM A370-07a. A370-07a Standard Test Methods and Definitions for Mechanical Testing of Steel Products. ASTM; 2007.