RESEARCH AND APPLICATION OF DOUBLE-CORE BUCKLING RESTRAINED BRACES IN TAIWAN

Transcription

1 13 th World Conference on Earthquake Engineering Vancouver, B.C., Canada August 1-6, 24 Paper No RESEARCH AND APPLICATION OF DOUBLE-CORE BUCKLING RESTRAINED BRACES IN TAIWAN Keh-Chyuan TSAI 1, Jiun-Wei LAI 2, Yean-Chih HWANG 3, Shen-Ling LIN 4 and Chong-Hsing WENG 5 SUMMARY In recent years, experimental and analytical researches have been conducted in NTU to investigate the effectiveness of various kinds of unbonding material, the brace-to-gusset connection details, analytical models and the design procedures for the Buckling restrained braces (BRBs). Tests conducted on more than 5 specimens confirm that the proposed BRBs employing double steel cores encased in twin steel tubes and infill concrete can stably sustain severe inelastic cyclic axial strain reversals. The double-cored BRBs can be conveniently connected to the gusset plate as in the traditional double-tee brace to gusset plate connections. Experiments include cyclic tests of BRB components and large scale single bay V- shaped buckling restrained braced frames (BRBFs). Test results indicate that the 2 mm thick silicon rubber sheets are effective in minimizing the difference between the cyclic peak compressive and tensile forces of the BRBs. The three V-shaped BRBF tests confirm that the steel brace core strain demands can be satisfactorily predicted from the story drift demand by geometry and incorporating the inelastic core length ratio. Test results also reveal that at a large story drift, the tensile strain in the tension brace was always greater than the compressive strain in the compression brace. This paper introduces all metallic BRB designs, non-destructive evaluation techniques for damage detections of BRBs, and the pseudo dynamic tests conducted on a full scale 3-story 3-bay CFT/BRB composite frame in NCREE in a Taiwan- US-Japan cooperative research project. The paper concludes with the applications of the proposed BRBs in seismic retrofit as well as new construction projects in Taiwan. INTRODUCTION Buckling restrained braced frame (BRBF) has been evolved into a very effective system for severe seismic applications (Watanabe et al. 1988). Buckling restrained braces (BRBs) or Unbonded Braces (UBs) commonly found are made from encasing a core steel cross-shape or flat bar member into a steel tube and confined by infill concrete (Fig. 1). The steel core member is designed to resists the axial forces with a full tension or compression yield capacity without the local or global flexural buckling failure. When the brace is subjected to compressions, an unbonding material placed between the core member and the infill 1 Director, NCREE, Taipei, Taiwan. kctasi@ncree.gov.tw 2 Research Assistant, NCREE, Taipei, Taiwan. jwlai@ncree.gov.tw 3 Senior Aircraftman, Air Force, Taiwan. huangyz@mail.apol.com.tw 4 Reserve Officer, Army, Taipei, Taiwan. iplay@ms31.hinet.net 5 Project Engineer, J.J. Pan & Partners, Taipei, Taiwan. r @ms89.ntu.edu.tw

2 concrete is required to reduce the friction. Thus, a BRB or an UB basically consists of three components, including steel core member, buckling restraining part and the unbonding material. Figure 2 illustrates the typical cross sections of the BRBs or UBs proposed by various researchers. Most of these BRBs are proprietary, but their concepts are essentially similar. It can be found that the cross section of the steel core member is usually bi-axially symmetric, can be a cruciform, an H or a flat bar shape. The buckling restraining part can be constructed from mortar filled in the tube, reinforced concrete, reinforced concrete covered with FRP or all-metallic steel tubes. It has been found that a single cored brace to the gusset connection is typically a butt joint using several splice plates and two set of connecting bolts as illustrated in Fig. 3. In order to reduce the length and the number of bolts in the brace-to-gusset connection, the double-cored buckling restrained braces (DCBRBs) as illustrated in Fig. 4 have been developed (Lai and Tsai 21) and extensively tested (Lai and Tsai 21, Huang and Tsai 22, Weng and Tsai 22, Lin and Tsai 23) in National Taiwan University (NTU) and Taiwan National Center for Research on Earthquake Engineering (NCREE) in the past few years. The proposed BRB members can be conveniently connected to the gusset plate in the same manner as that in the traditional double-tee brace to gusset plate connections (Fig. 5). These researches include the investigation of the effective unbonding material (Lai and Tsai 21), cyclic and fatigue performance of the DCBRBs employing mortar infilled in the double tubes for A36 (Huang and Tsai 22) and A572 GR5 (Lin and Tsai 23) steel cores. The design criteria of the brace end connections and the inter-connecting ties between the two tubes have also been established through experimental tests (Weng and Tsai 22). Tests have also confirmed that the selfcompact concrete, much more cost-effective than the cement mortar, is a satisfactory alternative to restrain the core steel member in the BRBs (Lin and Tsai 23). Core Steel Member Concrete (Mortar) Unbonding Material Steel Tube Unbonded Brace Fig.1 Schematic of BRBs L b /2 Unbonding Materials Fig.2 Typical types of buckling restrained braces Core Steel Member Concrete (Mortar) Tab Plate L b /2 Steel Tube (Buckling Restrained Part) Buckling Restrained Brace Fig. 3 Splice plate details for single-tubed BRB end connections Fig. 4 Schematic of double-tubed buckling restrained brace Fig. 5 Double-tee to gusset connection details for double-cored BRB end join

3 RESEARCH ON DOUBLE-CORED BRBS Effects of Various Unbonding Material In order to find out which kind of material possessing satisfactory unbonding effects, a total of ten BRBs (identical in core cross sectional area as shown in Fig. 6 but varying in the unbonding material) were tested under cyclic increasing displacement at NTU (Lai and Tsai 21). Table 1 summarizes the unbonding material used for each specimen and the corresponding cyclic loading protocol. The standard one refers to the protocols similar to the one provided by SAC (1997). The test setup is given in Photo 1. The typical cyclic response of the specimen is given in Fig. 7. The test results of the ten specimens are summarized in Fig mm Table 1 Specimen schedule for C.L. & Sym. unbonding material tests 2 4 mm 8 mm Side View D mm 8 3 Side View D A A 24 2 Sec. AA mm B 1,8 1 2 Sec. BB B R= Sec. CC C 3 C PL 216 x 64 x 8 mm A572 Grade mm 8 Sym. 4 C. L. & Sym unit : mm Fig. 6 Details and geometry of steel core element for ten unbonding materials testing specimens Photo 1 Experimental setup for BRB tests Axial Force ( kn ) SR2-1 Gauge Length = 18 mm 1.5 Py -2 A net = 36 mm Axial Deformation ( mm ) Fig.7 Cyclic response of the specimen using 2-mm silicone rubber sheet unbonding material Py Axial Load Difference.1 32 Specimen (1) Unbonding Material (2) Unbonding Material Thickness (3) Loading History (4) AS-1 Asphalt Paint N. A. Standard VF-1 VK-1 Vinyl Sheet + Foaming Tape Vinyl Sheet + Kraft Tape 2 mm Standard 2 mm Standard R2-1 Rubber Sheet 2 mm Standard R5-1 Rubber Sheet 5 mm Standard SR1-1 SR2-1 SR2-2 SR2-3 SR5-1 Silicone Rubber Sheet Silicone Rubber Sheet Silicone Rubber Sheet Silicone Rubber Sheet Silicone Rubber Sheet Cycle Number 1 mm Standard 2 mm Standard 2 mm Low-Cycle Fatigue 2 mm Near-Fault 5 mm Standard AS-1 VK-1 ε =.15.4 R5-1 ε =.2 R2-1 ε =.25.3 SR1-1 SR5-1 SR2-1.2 SR2-2 Fig.8 Axial load difference under cyclic loading For the purpose of discussion, the axial load difference Γ is defined as: ( Cmax Tmax) Γ = (1) Tmax where C max and T max are the maximum compressive and tensile brace forces at a same absolute axial deformation level. Theoretically speaking, after the core member is yielded, the Poisson ratio ν =.5 can be applied in the following calculations. The volume of the yielding steel segment should remain constant, that is: A L = A L (2) %

4 where A o and L o correspond to the original core cross sectional area and length, respectively, while A and L correspond to those after the brace is deformed in either tension or compression. Therefore, it can be shown that the axial strain is: L A ε = 1 = 1 and A = A (1 ε ) (3) L A Thus, the ratio between the compressive and tensile brace forces for a given (absolute) strain level could be: ( Cmax Tmax ) A (1 + ε ) A (1 ε ) 2ε Γ = = = 2ε (4) Tmax A (1 ε ) 1 ε Equation 4 suggests that the Γ is about 4 % for an ε = 2 %. But the test results shows in Fig. 6 exhibit much higher Γ values (the maximum is about 3% for ε = 2 %). This should be due to the imperfect unbonding mechanism and a substantial friction developed between the steel core member and the buckling restraining part. It is found in Fig. 8 that the 2-mm thick silicone rubber sheet has the least axial load difference (about 1% at an axial strain of 2%) under the cyclic increasing displacements. Therefore, for the subsequent tests of the BRBs, unless it is otherwise noted, the 2-mm thick silicone rubber sheets have been adopted for the construction of most of the specimens. In Fig. 7, it should be noted that the strain hardening effects are evident, comparing with the tensile yield capacity computed from the tensile coupon strength. Key Mechanical Properties of the BRBs In order to properly confine the BRB s inelastic deformations inside the restraining tube, the cross sectional area (A c ) of the energy dissipation core segment (L c ) is smaller than that of the end joint regions (L j ). Schematic configuration of a DCBRB in the frame is illustrated in Fig. 9, in which L c and L wp represent the core length and the work-point to work-point length, respectively. Between the end and the core segment, a transition region as illustrated in Fig. 1 can be devices. It is confirmed by tests (Huang and Tsai 22) that the effective stiffness, K e of the BRB considering the variation of cross sectional area along the length of the brace, can be accurately predicted by: EA j Ac At K e = (5) A A L + 2 A A L + 2A A L j t c c t j c j t Area = At Total Length = Lt Area = Ac Total Length = Lc Area = Aj Total Length = Lj L c L wp Lj/2 Lt/2 Lc Lt/2 Lj/2 W.P. Fig. 9 Dimensions of core length and work point to work Fig. 1 Profile of steel core member in the double tube BRBs point length The relationships between the brace overall strain (ε wp ) and the inter-story drift θ can be approximated as: ε wp = θ sin 2φ / 2 (6) where φ is the angle between the brace and the horizontal beam as illustrated in Fig. 11. The strain-to-drift ratio versus the beam angle φ relationships given in Eq. 6 are plotted in Fig. 12. If the ratio of the core length and the work-point to work-point dimension is: α= L c / L wp (7) W.P.

5 ~ θ δ θ θ = H { ~cos H ε θ wp.3 1 ( radian ).2.1 φ=45 ο φ (degree) δ δ θ= H φ H Experiment Analysis UB735C 1.5P y -1 Gauge Length = 182 mm Axial Deformation ( mm ) Fig. 11 Schematic of brace Fig. 12 Brace strain to story Fig. 13 Experimental and analytical deformation versus inter-story drift ratio versus brace angle responses of a typical double plate drift angle relationship relationship BRBs Thus, assuming the strain outside the core segment is negligible, the BRB core inelastic upper bound strain ε c can be express as: ε c ε wp / α (8) From Eqs. 6 through 8, it can be found that if the inter-story drift demand is.2 radian, then the peak core strain would be close to.2 for a BRB having a length aspect ratio α =.5 and oriented in a 45 degree angle. The cyclic responses of a typical DCBRB constructed with two A36 flat steel cores (Fig. 4) using the 2-mm silicone rubber unbonding sheets are shown in Fig. 13. It is evident in Fig. 13 that a strain hardening factor of about 1.5 is appropriate in estimating the peak tensile strength for A36 core member. In addition, it appears that the following equation should be applied in estimating the maximum compressive strength possibly developed in a BRB: P max = Ω Ωh β P y (9) where P y = A c F y is the nominal yield strength of the core section, Ω and Ω h take into account the possible material over-strength and strain hardening factors of the core steel, respectively. In addition, the bonding factor β represents the imperfect unbonding, the fact that the peak compressive strength is somewhat greater than the peak tensile strength observed during the large deformation cycles. Therefore, the required stiffness of the steel casing in order to prevent the BRB from a global flexural buckling is (Watanabe et al. 1988): 2 Pmax ( kl) I tube FS (1) 2 π E It is noted from Fig. 13 that for a properly fabricated BRB, a bonding factor β of at least 1.1 is appropriate if the peak cyclic core strain demand is no greater than.2. Figure 12 also indicates that the experimental responses of the BRB can be accurately represented by using the two-surface plastic hardening material model implemented for the truss element in a general purpose frame response analysis computer program (Tsai and Chang 21). Figure 14 further confirms that the same specimen, after subjecting the SEAOC/AISC (21) standard loading protocols, sustained a total of 262 cycles of large fatigue strain before fracture. As the flexural buckling of a single-cored BRB member under large compress strains could occur (Tsai et al. 22) at a section where the steel tube terminates (Fig 3), it is recommended that the following stability criterion be met for connection details shown in Figs. 3 and 5: P e_trans 2 π EI = P (11) trans 2 b ( kl ) max where P max is given in Eq. 9 and EI trans is the flexural stiffness of the core member at a section near the end of the steel tube. As noted earlier, the double-cored BRBs can be conveniently connected to the gusset plate as shown in Fig. 5. As a result, the connection length is reduced as only one set of connecting bolts is required at each brace end. The connection length can be further reduced if a welded detail directly P y

6 attaching the three edges of each tee to the gusset plate is adopted. An application example of the welded brace end joint in a SRC building will be given later in this article. 15 Steel Tube 1 1.5P y P y Axial Force ( kn ) th Cycle Bolt Holes Tab Plate Locations -1 Component Fatigue Axial Deformation ( mm ) Fig. 14 Fatigue responses of the Specimen UB735 Steel Tube Fig. 15 Schematic of welded tube-to-tube tie connections Fig. 16 Several possible tie connection details Slip Resistant Bolted Connection Details in Brace End Double Tee to Gusset Joints In 22, six BRBs featuring double-plate core with double tee end details were tested at NTU (Weng and Tsai 22) incorporating two different levels of roughness at the tee-to-gusset contacting surfaces. A set of cyclically increasing forces and displacement was applied to find out the slip load of the brace end bolted joints. Then two addition sets of similar cyclically increasing forces and displacements but reduced in magnitude subsequently were applied to confirm the cyclically degrading of the slip capacity. Test results suggest that using 1.5 times the strength of the bolt for design can prevent slipping of the connection under cyclic loading reversals. That is: R str = N N 1.5 F A ) (12) s b ( v b where R str is the slip-resistant strength of the connection, N s is the number of slip surface, N b is number of bolts, F v is the nominal shear strength of the bolt, and A b is the nominal tension area of the bolt. This research also investigated the longitudinal and transverse strain distributions in the web of the tee at the brace-to-gusset joints and concluded with the design recommendations for the brace end connection details. Tube-to-Tube Tie Connection Designs for Double-Tubed BRBs Since the proposed double-cored BRB consists of two independent units of bracing, tie connections between two units can be continuous or properly spaced. These ties can be made by welded bars (tab plates) as illustrated in Fig. 15 or various possible bolted details as illustrated in Fig. 16. It is found that by using the elastic stability theory (Timoshinko and Gere 1961), the required strength P req and stiffness β id of the tie connection can be derived (Weng and Tsai 22): 3 P B σ ( y P P ) ( ) L 2 E e req = cr + tube (13) 9 Pmax E e βid = ( Pcr ) (1 + ) 2 L 2 B σ (14) tube y where L tube is the length of the buckling restraining tube, P max is the maximum axial force suggested in Eq. 9 for the DCBRB, P cr is the critical eccentric load of the single tube derived from Secant formula (Timoshinko and Gere 1961), B is the width of short side of the rectangular tube, E is Young s modulus of steel, σ y is yield stress of steel tube, e is the eccentricity of load measured from neutral axis of the single tube. For example, if the design strength P y of the DCBRB is equal to 196 kn, and estimated maximum axial force P max = 485 kn, then according to the stiffness requirements noted in Eq. 14, the size of the each buckling restrained tube is mm, with e = 32 mm, L tube = 454 mm, B = 15 mm, P cr can be calculated from the Secant formula as 81 kn. Based on Eqs. 13 and 14, the required strength P req and stiffness β id of the tie connections at every 1/3 of the tube length are 38 kn and 36 kn/mm, respectively.

7 Tests confirm that the proposed strength and stiffness requirements developed for the tie connection elements between the twin tubes can be conveniently applied in the design and construction of the doubletubed BRBs subjected to large inelastic strain reversals. Tests on Single Story V-shaped BRB Frames In order to assess the performance of the double-t to gusset connection details, three large scale single bay V-shaped buckling restrained braced frames (BRBFs) constructed with the proposed BRBs have been tested in NCREE (Huang and Tsai 22). The objectives of the study include: 1) investigating the experimental and analytical responses of the single bay V-shaped BRBFs each constructed with two BRBs in three different length aspect ratios, 2) investigating the steel BRB core strain versus inter-story drift relationships, and 3) providing guidelines for the analysis and design of BRBF for severe seismic applications. The schedule of three frame specimen and one BRB component is given in Table 2. The profile of all the BRBs is shown in Fig. 4. The frame elevations of these specimens are given in Figs.17 and 18. The experimental setup for the frame test is given in Photo 2, while the BRB to gusset connection details can be found in Photo 3. The loading protocol for the frame tests and the components are illustrated in Figs. 19 and 2, respectively. Table 2 Specimen schedule for BRB frame tests Frame (1) BRB Specimen (2) Yield Strength (kn) (3) Total Bolts (F1T 24mm) (4) α (5) Peak Drift (% rad) (6) Cycles for Fatigue Test (7) NSYF SYMF SYMSCF NSY88W NSY59E SYM735W SYM735E SYM735SCW SYM735SCE UB735C until Breaking Fig. 17 Elevation of the Specimen NSYMF Fig. 18 Elevation of the Specimen SYMF Photo 2 Experimental setup for BRB frame test (Specimen SYMF) Photo 3 Brace end double tee to gusset connection details The cyclic and the fatigue performance of the UB735 are shown in Figs. 13 and 14, respectively. The experimental cyclic force versus deformation responses (Fig.13) of the BRB component exhibit stable energy dissipation characteristics. Fig. 13 also shows that experimental BRB responses can be accurately predicted using an inelastic truss element incorporating the two-surface plasticity model implemented in a general purpose finite-element computer program (Chang and Tsai 21). Fig. 14 indicates that same

8 θ Story Drift (% radian) N Displacement between two beams θ Story Drift (% radian) kn 49 kn 735 kn Force Control Fig. 19 Cyclic loading protocol Fig. 2 Cyclic loading protocol for UB735 specimen after going through the inelastic excursions shown in Fig. 13, sustained a total of 262 cycles of fatigue strain on the order of.125 before fracture. Fig. 17 shows that the arrangement of BRBs can be configured un-symmetrically in order to accommodate architectural function needs, such as door or window openings. In this manner, the cross sectional area and the length of the energy dissipation segments of a pair of braces can be specifically tailored to avoid the potential unbalanced vertical load resultants while reach yielding at pretty much the same time. As shown in Figs. 21 through 23, the cyclic or fatigue performances of the frame systems and the BRBs indicate the proposed BRB components and framing system possess extremely stable characteristics. In addition, as shown in Figs. 24 and 25, the inelastic finite element models accurately predicts the frames and the BRBs experimental lateral force versus story drift responses for all test frames (only shown for Specimen NSYF herein). The three V- shaped BRBF tests confirm in Fig. 26 that the steel brace core strain demands can be satisfactorily predicted from the story drift demands by geometry (Eq. 6) and incorporating the ratio of the work pointto-work point dimension to the inelastic core length (Eqs. 7 and 8). Test results in Figs. 26 and 27 also reveal that at a large story drift, the tensile strain in the tension brace was always greater than the compressive strain in the compression brace. This phenomenon is more pronounced as the inter-story drifts increase. This somewhat suggests that the peak compression and tension forces have a tendency to self-equilibrate as story deformation increases and reach a reduced unbalanced vertical force components resisted by the horizontal beam member NSYF Cyclic SYMF Cyclic SYMSCF Cyclic Fig. 21 Total lateral force versus story drift relationships 2 N Peak Actuator Displacement (mm) All Metallic and Detachable BRBs As noted above, single- or double-cored BRBs properly made from cement mortar or concrete filled in the steel casing possess satisfactory seismic resistant characteristics. However, as shown in Fig. 28, it does not allow visual inspecting the damage state of the inner core after an earthquake. In order to simplify the fabrication of the BRBs made from using infill mortar and unbonding coating, the all-metallic welded BRB as shown in Fig. 29 or the fully detachable BRBs illustrated in Fig. 3 have been developed and extensively tested in NTU (Tsai and Lin 23). The detachable features shown in Fig. 3

9 3 2 NSYF Cyclic 3 2 SYMF Cyclic 3 2 SYMSCF Cyclic Brace Shear ( kn ) 1-1 Brace Shear ( kn ) 1-1 Brace Shear ( kn ) NSYF Fatigue Fig. 22 Total brace force versus story drift relationships SYMSCF Fatigue SYMF Fatigue Fig. 23 Fatigue performance of three BRB frames without failure NSYF Experiment Analysis Fig. 24 Experimental versus analytical frame shear versus lateral drift responses of Specimen NSYF Core Strain NSYF East ε core =1. 1θ Compression Tension Theory Story Drift (radian) Core Strain SYMF Eest Brace Shear ( kn ) NSYF Experiment Analysis Fig. 25 Experimental and analytical total brace shear versus story drift relationships in Specimen NSYF ε core = 1. 22θ Compression Tension Theory Story Drift (radian) Core Strain SYMSCF East ε = 2. core 44θ Compression Tension Theory Story Drift (radian) Fig. 26 Experimental versus predicted core strain versus inter-story drift relationships for three frames provide the possibility of disassembling the BRBs for inspection after an earthquake. In order to allow the extension and contraction of two ends of a BRB, a stopper shown in Fig. 31 to lock into the restraining concrete or other buckling restraining part has been adopted to prevent the buckling restrainer from slipping off. Instead of applying the unbonding coating, a 1-mm or 2-mm small gap has been experimented between the surface of the inner core and external restraining tubes. A total of thirteen detachable BRB specimens using A572 GR5 steel plates were fabricated and tested using the standard test protocol given in Fig. 32. Test results indicate that a 1-mm or 2-mm gap between the inner core and tubes is most effective in minimizing the difference between the peak BRB compressive and axial tensile force responses as illustrated in Fig. 33. After applying a standard loading protocol shown in Fig. 32, the specimen was never fractured. The same specimen was then subjected to a constant fatigue strains until fractured. Test results suggest that the all metallic and detachable BRBs can stably sustain severe cyclic increasing and constant fatigue inelastic axial strain reversals. As the specified constant strain varied from

10 specimen to specimen, therefore, the number of fatigues cycles versus the fatigue strain relationships can be constructed as shown in Fig. 34. In addition, the cumulative plastic deformation (in terms of y ) is constructed and given in Fig. 35. Tests confirm that a typical fully detachable BRB exhibit stable and excellent hysterestic behavior as any other BRBs or UBs (restrained by mortar and steel casing) tested before. It is noted that the strain hardening factor for a BRB made from A572 GR5 is about 1.25 for a peak core strain less than.25. In this series of tests, it is found that the self compact concrete (SCC) with a compress strength of 56 MPa is a cost-effective alternative to the cement mortar for the infill in a steel casing. Core Strain SYMF Drift(-) West Brace(+) Drift(-) Eest Brace(-) Drift(+) West Brace(-) Drift(+) Eest Brace(+) Core Strain SYMSCF Drift(-) West Brace(+) Drift(-) Eest Brace(-) Drift(+) West Brace(-) Drift(+) Eest Brace(+) Story Drift (radian) Story Drift (radian) Fig. 27 Tensile strain versus compressive strain at peak inter-story drift for Specimens SYMF and SYMSCF connection transition segment segment yielding segment A B partition confining-bolt block angle unbonding material buckling restrained mechanism Fig. 28 Steel cores are confined in the steel tubes without exposing the energy dissipating segments Core Steel Member Steel Tube (Buckling Restrained Mechanism) Tab Plate Buckling Restrained Brace Fig. 29 Schematic of welded all metallic BRBs A confining-bolt core segment B steel buckling-restrained mechanism joining-bolt plate tube core joining-bolt Sect. A-A Sect. B-B core tube plate Fig. 3 Schematic of the bolted detachable BRBs θ Drift Angle (% radian) 1.5D bm D bm =1.5% 3 R D/d=1.2, t/w=.2, R/W=.4 D d W t=.1d, W=.5d, R=.2d t stopper detail Fig. 31 Details of the stopper D bm.5d bm.5p y Force Control N N Number of Cycles Fig. 32 Standard loading protocol (SEAOC/AISC 21) DAMAGE INSPECTION AND NON-DESTRUCTIVE TESTING OF BRBS As the energy dissipation segment of a conventional BRB is generally covered by the steel casing, direct damage detection may not be so straight forward. In order to inspect the BRBs after severe services, nondestructive test (NDT) techniques have been applied on more than 2 BRBs before, during and after the load tests (Tsai and Lin 23). During various stages of each one specimen s load test, it is done by simultaneously hitting the BRB at one end and measuring the stress wave at the opposite end as illustrated

11 in Figs. 36 and 37. The NDT tests were conducted by hitting and observing the variation of stress wave form measurements collected from the BRB specimens. Receiver No. 2 shown in Fig. 37 was CPD Axial Force ( kn ) Strain=(U3+U4+U5+U6)/2/L c 1.25 Py Py -2 HG2 Standard Core Strain 1, 1, Axial Force ( kn ) Strain=(U3+U4+U5+U6)/2/L c 1.25 Py Fig. 33 Force versus strain relationships for typical A572 GR5 BRBs CPD=1.443(ε c ) R 2 =.926 Amplifier 1 Py 28th Cycle 37th Cycle HG2 Fatigue (.15) Core Strain Oscillograph Amplifier 2 Number of Loading Cycles N(cycle) N=.7(ε c )-2.25 R 2 =.972 Double Slide series All-Metallic series Core Plastic Strain Range ε c Fig. 34 Number of fatigue strain cycles versus core strain relationship (A572 GR5 BRBs) S Impact Receiver 2 Receiver 1 L L R R N Double Slide series All-Metallic series Core Plastic Strain Range ε c Fig. 35 Cumulative plastic deformaiton ( y ) versus fatigue strain relationships SCC_P_L_L Step 5 SCC_P_L_R Receiver 1 Receiver 2 Impact Source Fig. 36 Schematic setup of impact and stress wave measurement Step 5 SCC_P_R_R Fig. 37 Locations of impact and measuring points Step 5 SCC_P_R_L Step 5 Wave Shape Step 4 Step 3 Step 2 Wave Shape Step 4 Step 3 Step 2 Wave Shape Step 4 Step 3 Step 2 Wave Shape Step 4 Step 3 Step 2 Step 1 Step 1 Step 1 Step Time (1-6 sec) Time (1-6 sec) Time (1-6 sec) Time (1-6 sec) Fig. 38 Stress wave forms recorded from the impact tests on a BRB using the self-compact concrete or SCC (P stands for P wave, L-R denotes the location of the Receivers 1 and 2) needed to normalize the impact force before comparing the impact and the response at the two BRB ends, respectively. A typical set of stress wave form is shown in Fig. 38, where the hitting/measuring Steps 1 to 3 were conducted, respectively, when specimen was on the floor, after connecting to the universal testing machine and after applying the standard loading protocol. Thus, the wave forms shown in Steps 1 to 3 in Fig. 38 suggest that the BRB had not been fracture. Step 4 was conducted after the same specimen had been fractured under the fatigue stain, and Step 5 was done after the specimen was removed from the load frame and brought down to the floor. It is evident in Fig. 38 that the stress wave forms are much calm after the fatigue fracture test. These results seem to suggest that the proposed NDT technique can be a simple way to detect whether the inner core is fractured or not. In addition, it is found in this study that the pressure wave (P-wave) measurements are more satisfactory than those of the shear wave (S-wave) for the stress wave NDT of BRBs.

12 PSEUDO DYNAMIC TESTS OF A FULL SCALE 3 STORY CFT/BRB COMPOSITE FRAME Through international collaboration between researchers (Goel and Tsai 24) in Taiwan, Japan, and the United States, a full-scale 3-story 3-bay RC column and steel beam RCS composite moment frame was tested in NCREE in October 22 (Chen et al. 24) using networked pseudo dynamic test procedures. In October of 23, a full-scale 3-story 3-bay composite braced frame (CFT/BRBF) consisting of CFT columns and BRB members has been tested similar. The 3-story prototype structure is designed for a highly seismic location either in Taiwan or United States using displacement based seismic design procedure (Tsai et al. 23). The typical bay width is 7m and typical story height is 4m. The total height of the frame, including the footing, is about 13m. The 215mm wide concrete slab was adopted to develop the composite action of the beams. Measuring 12 meters tall and 21 meters long, the specimen is among the largest frame tests of its type ever conducted (Photo 4). Three different types of moment connections, namely through beam, external diaphragm and bolted end plate types, varying from the first floor to the third floor were fabricated for the exterior beam-to-column connections. Three types of BRBs, including the single core, double-cored and the all metallic BRBs, were adopted in the three different stories. In particular, two single-cored Unbonded Braces, each consisting of a steel flat plate in the core, were donated by Nippon Steel Company and installed in the second story. The two BRBs installed in the third story are double-cored constructed using cement mortar infilled in two rectangular tubes (Tsai et al. 23). The frame has been tested using the pseudo-dynamic test procedures applying input ground motions obtained from the 1999 Chi-Chi and 1989 Loma Prieta earthquakes, scaled to represent 5%, 1% (DBE), and 2% (MCE) in 5 years seismic hazard levels. Following the pseudo-dynamic tests, since none of the brace was fractured, quasi-static loads were applied to cyclically push the frame to large inter-story drifts up to the failure of the braces. Being the largest and most realistic composite CFT/BRB frame ever tested in a laboratory, the test provides a unique data set to verify both computer simulation models and seismic performance of CFT/BRB frames. These two large scale frame tests have also provided great opportunities to explore international collaboration and data archiving envisioned for the Networked Earthquake Engineering Simulation (NEES) or the Internet-based Simulations for Earthquake Engineering (ISEE) research programs launched recently in USA and Taiwan, respectively. The analytical predictions were broadcasted along with the real time experimental results during each test as shown in Fig. 39. Additional information can be found in the related web sites ( and Fig. 39 On-line data broadcasting of analytical prediction and experimental results during and after the tests ( EXAMPLE APPLICATIONS OF DOUBLE-CORED BRBS IN TAIWAN Several new and retrofit projects in Taiwan have selected the double tubed BRBs as the energy dissipation element to improve the buildings seismic performance. A 1-story gymnasium at Chinese Culture University in Taipei has been proposed and is being constructed. The lateral force resisting structural system consists of mega braced frames in the longitudinal direction and build-up truss moment resisting frames in the transverse direction as shown in Photo 5. A total of 96 pieces of double-tube BRBs is being

13 installed in the three-dimensional steel frame. In this project, the maximum length of the double-cored BRB is about 11 meter long and the A572 GR5 steel core plate size is 45 mm 26 mm, Photo 4 Experimental setup for pseudo dynamic tests of a 3-story 3-bay full scale CFT/BRB composite frame each inside a concrete-filled 5mm 25mm 15mm built up steel box casing. The peak axial tensile and compressive strengths could reach almost 1, kn. In order to confirm the cyclic performance of the large size double-cored BRBs, four full scale BRB components have been tested at NCREE (Photo 6) using a loading protocol similar to that given in Fig. 32. Test results given in Figs. 4 and 41 confirm that the full scale large size double-tubed BRBs can sustain a large number of inelastic cyclic increasing and constant fatigue strain reversals before failure. In addition, the cumulative inelastic deformation exceeds 14, except for specimen U2A ( See Table 3 ) due to additional rotation at brace ends induced by test setup, well meets the SEAOC/AISC prescribed requirements (21). Photo 7 shows the construction site at Chinese Culture University in February this year. Photo 8 and 9 shows the construction site and BRBF details of a 14-story TV Station in Taipei. A total of 96 pieces of double-tubed BRBs are being installed in the opposite perimeter bays along the transverse direction of this SRC structure. It is shown in Photo 8 that the welded details have been adopted for the brace end connections effectively reducing the size of the double tee to gusset plate joint. Shee-Hwa United World Tower ( Photo 1 ) is a 46-story office building in Taichung, designed before 1999, being seismically upgraded to accommodate an increased seismic hazard level. A total of 8 pieces of double-tubed BRBs are being installed in the two opposite perimeter bays along the longitudinal direction of the frame. Photo 11 shows the construction site and BRB detail of Shee - Hwa United World Tower. 1 Core Strain (mm / mm) U1A - Standard 1.25 F y Axial Force (kn) 5-5 F y -F y F y Photo 5 Three dimension model of the gymnasium Photo 6 Experimental set up of full scale BRB tests Axial Deformation (mm) Fig. 4 Cyclic force versus deformation relationships

14 Core Strain (mm / mm) U1A - Fatigue 1.25 F y Axial Force (kn) cycles F y -F y F y Photo 7 Construction Site of Gymnasium Table 3 Cumulative Plastic Deformation of Specimens Protocol Type Standard Low Cycle Fatigue Total Cycle U5B U2A T69-2 U1A Level ±1. ± D bm ±1. D bm ±1.5 D bm ±2. D bm CPD ±1. D bm CPD CPD N. A N. A. N. A. N. A N. A. N. A Axial Deformation (mm) Fig. 41 Fatigue force versus strain relationships Photo 8 Construction Site of Tzu-Chi TV Station Photo 9 BRBF detail of Tzu-Chi TV Station Photo 1 The 3D Model of Shee-Hwa United World Tower Photo 11 Construction Site and BRB details of Shee- Hwa United World Tower ACKNOWLEDGEMENTS Taiwan National Science Council is gratefully acknowledged for the financial supports provided for research on buckling restrained braces. The technical supports provided by the laboratory at Taiwan National Center for Research on Earthquake Engineering are gratefully acknowledged. During these studies, technical information provided by Nippon Steel Company of Japan is deeply appreciated. Many graduate students made significant contributions in their master thesis works on BRB research in the

15 Structural Engineering Division of Department of Civil Engineering at NTU. These students include Jiun- Wei Lai, Yean-Chih Huang, Chong-Hsing Weng, Shen-Ling Lin and Po-Chien Hsiao. Without their efforts, these research results would not be possible. Valuable suggestions provided by many researchers from Taiwan, Japan and US during the full scale CFT/BRB frame tests are gratefully acknowledged. REFERENCES 1. Chang, L.C. and Tsai, K.C. (21), The Platform and Visualization of Inelastic Structural Analysis of 2D Systems PISA2D and VISA2D, Report No. CEER/R9-8, Center for Earthquake Engineering Research, National Taiwan University. (in Chinese) 2. Chen, C.H., Hsiao, P.C., Lai, J.W., Lin, M.L., Weng, Y.T. and Tsai K.C. (24), Pseudo-Dynamic Test of a Full-Scale CFT/BRB Frame: Part 2- Construction and Testing. Proceedings of 13th World Conference on Earthquake Engineering, Vancouver, B.C., Canada, Paper No. 2175, August Lin, M.L., Weng, Y.T., Tsai, K.C., Hsiao, P.C., Chen, C.H. and Lai, J.W. (24), Pseudo-Dynamic Test of a Full-scale CFT/BRB Frame: Part 3 - Analysis and Performance Evaluation. Proceedings of 13th World Conference on Earthquake Engineering, Vancouver, B.C., Canada, Paper No. 2173, August Chen, C.H., Lai, W.C., Cordova, P., Deierlein, G.C. and Tsai, K.C. (24), Pseudo Dynamic Tests of a Full Scale RCS Frame: Part 1: Design, Construction, Testing, Structures Congress, ASCE, Nashville. 5. Huang, Y.C. and Tsai, K.C., (22), Experimental Responses of Large Scale Buckling Restrained Brace Frames, Report No. CEER/R91-3, Center for Earthquake Engineering Research, National Taiwan University. (in Chinese) 6. Lai, J.W., Tsai, K.C. (21), A Study of Buckling Restrained Brace Frames, Report No. CEER/R9-7, Center for Earthquake Engineering Research, National Taiwan University. (in Chinese) 7. S.L. Lin and Tsai, K.C. (23), A Study of All Metallic and Detachable Buckling Brarces, (in press) Center for Earthquake Engineering Research, National Taiwan University. (in Chinese) 8. SAC (1997), Protocol for Fabrication, Inspection, Testing, and Documentation of Beam-Column Connection Tests and Other Experimental Specimens. Appendix E: Loading Protocol for Stepwise Increasing Cyclic Tests, SAC Background Document SAC/BD-97/2. 9. SEAOC/AISC (21), Recommended provisions for buckling-restrained braced frame (draft), SEAOC and AISC. 1. Tsai, K.C., Huang, Y.C., Weng, C.S., Shirai, T. and Nakamura, H. (22), Experimental Tests of Large Scale Buckling Restrained Braces and Frames, Proceedings, Passive Control Symposium, Tokyo Institute of Technology, Tokyo. 11. Timoshenko, S.P., and Gere, J.M. (1961), Theory of Elastic Stability, 2nd Ed., McGraw-Hill Book Company. Inc,New York. 12. Goel, C.S., Tsai, K.C. (24), overview of international cooperative research on seismic performance of composite and hybrid structures, Proceedings, Structures Congress, ASCE, Nashville. 13. Tsai K.C., Weng, Y.T., Lin, M.L., Chen, C.H. Chen, Lai, J.W. and Hsiao, P.C. (23), Pseudo Dynamic Tests of a Full Scale CFT-BRB Composite Frame: Displacement Based Seismic Design and Performance Evaluations, Proceedings, International Workshop on Steel and Concrete Composite Constructions, National Center for Research on Earthquake Engineering, Taipei, Taiwan. 14. Tsai, K.C., Lai, J.W., Chen, C.H., Hsiao, P.C., Weng, Y.T., and P.C. Hsiao (24), Pseudo Dynamic Tests of a Full Scale CFT/BRB Composite Frame, Proceedings, Structures Congress, ASCE, Nashville. 15. Tsai, K.C., Weng, Y.T., Lin, S.L., and Goel, subhash (24), Pseudo-Dynamic Test of a Full-Scale CFT/BRB Frame: Part 1- Performance Based Specimen Design. Proceedings of 13th World Conference on Earthquake Engineering, Vancouver, B.C., Canada, Paper No. 75, August Watanabe, A., Hitomoi, Y., Saeki, E., Wada, A., and Fujimoto, M. (1988), Properties of Brace Encased in Buckling-Restraining Concrete and Steel Tube, Proceedings of Ninth World Conference on Earthquake Engineering, August 2-9, Tokyo-Kyoto, Japan.(Vol. IV-719~724) 17. Weng, C.H. and Tsai, K.C., (22), Experimental Responses of Double-Tube Unbonded Brace Elements and Connections, Report No. CEER/R91-2, Center for Earthquake Engineering Research, National Taiwan University. (in Chinese)