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THE STRUCTURAL DESIGN OF TALL AND SPECIAL BUILDINGS Struct. Design Tall Spec. Build. 19, 77–93 (2010) Published online 30 October 2009 in Wiley Interscience (www.interscience.wiley.com). DOI: 10.1002/tal.543 CASE STUDY OF A 40-STOREY BUCKLING-RESTRAINED BRACED FRAME BUILDING LOCATED IN LOS ANGELES ANINDYA DUTTA* AND RONALD O. HAMBURGER Simpson Gumpertz & Heger, San Francisco, California, USA SUMMARY Simpson Gumpertz & Heger has prepared two prototypical designs for a 40-storey buckling-restrained braced steel-framed office building located at a generic site in Los Angeles, CA. One of these designs conforms in all respects, except height limits with the design criteria contained in the 2007 California Building Code and ASCE 7.05 Standard for Minimum Design Loads for Buildings and Other Structures. The second design has been conducted using a performance-based approach generally based on the criteria contained in guidelines published by the Los Angeles Tall Buildings Council. The performance-based design incorporates fewer bays of bracing and lighter members than the code-based design, but is intended to provide performance at least equivalent to that anticipated for code-designed buildings. The purpose of this work was to permit study of the performance capability of buildings designed to alternative criteria. This work was performed in support and under funding provided by the Pacific Earthquake Engineering Research Center’s Tall Buildings Initiative. Copyright © 2009 John Wiley & Sons, Ltd. 1. OBJECTIVE AND SCOPE OF WORK Simpson Gumpertz & Heger, Inc. developed two alternative designs of a 40-storey steel building on behalf of the Pacific Earthquake Engineering Research Center (PEER) under its Tall Buildings Initiative (TBI). The objective of the TBI is to develop recommended performance-based design criteria for tall buildings as an alternative to the criteria contained in present building codes and standards. These alternative performance-based criteria are intended to provide performance that is at least equivalent to that intended for buildings designed in conformance with the code. The buildings presented in this paper, together with designs using other structural systems developed by other designers, will be used by PEER to assess the cost and performance capability of buildings designed using alternative design approaches and criteria. This information will guide the development of the PEER recommendations. We developed the designs to a schematic level. We sized the gravity load system, including floors, systems and columns considering blanket superimposed and live loads, but neglecting miscellaneous openings, cladding supports and similar information that typically becomes available later in the design process. Seismic analysis, including linear response spectrum analysis for the code-based design and nonlinear response history analysis for the performance-based design, was performed, and members of the seismic force-resisting system were proportioned as required to meet the respective criteria. We designed the foundations to a sufficient level to determine mat thickness, but did not determine reinforcing. Detailing of the structure was not performed. * Correspondence to: Anindya Dutta, Simpson Gumpertz & Herger, The Landmark at One Market, Suite 600, San Francisco, CA94105. E-mail: [email protected] Copyright © 2009 John Wiley & Sons, Ltd. 78 A. DUTTA AND R. O. HAMBURGER In parallel with our work, two other firms performed similar designs of buildings using reinforced concrete structural systems. All buildings were assumed to be located on the same generic site in downtown Los Angeles. PEER provided site criteria for us in our design including site class, permissible bearing pressures, spectral response ordinates and ground motion acceleration histories. Our scope of work included the following: (1) Meet with representatives of PEER and other design consultants engaged in the TBI project to develop the general criteria for the designs. (2) Develop a design, to a schematic level, for an essentially code-conforming steel-framed building. (3) Develop a design, to a schematic level, for a performance-based steel-framed building, of similar height and footprint, using the procedures contained in design recommendations prepared by the Los Angeles Tall Buildings Council. (4) Prepare schematic-level drawings documenting the design. 2. BUILDING DESCRIPTION Both designs have typical above-grade floors comprised of 6 ¼ in. lightweight concrete fill on metal deck supported by composite steel framing with a foot print of 170 by 107 ft. Both designs have four basement levels with a foot print of 227 by 220 ft. Both structures have lateral force-resisting systems comprised of buckling-restrained braced frames without backup moment frames. The site class was assumed as class C. Spectral response ordinates were: Ss = 2·15 and S1 = 0·72. For the performance-based design, we used a suite of seven acceleration histories provided by PEER. 3. 3.1 CODE-BASED DESIGN Purpose This section provides a brief overview of the code-based design. As discussed earlier, we used the 2007 California Building Code to perform this design. 3.2 Design description This design was conducted in conformance with all the prescriptive provisions of the 2007 California Building Code and its referenced standards, except for the limitation on the building height. The California Building Code requires that buildings in excess of 160 ft and located on the site selected for this building incorporate a special moment-resisting frame capable of resisting at least 25% of the specified design seismic forces. This design did not incorporate such a frame. Figure 1 shows building plan views at various levels, and salient features of this design are described further below. 3.3 Gravity analysis and design We designed the building’s vertical load-resisting system using gravity loads that include a combination of structure self-weight and additional superimposed loads. The design criteria document provided to us by PEER specified the superimposed loads at various floors shown in Table 1. We combined the element self-weights with the superimposed loading from this table, and performed a gravity-load design using the RAM Structural System software, version 12.1. Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB 79 Figure 1. Plan view of the building at various floors Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 80 3.4 A. DUTTA AND R. O. HAMBURGER Lateral analysis Following design of the vertical load-carrying elements, we performed a lateral analysis of the building. We considered both wind and seismic forces, and designed each element for the most severe requirements. A brief description of the wind and seismic design is given below. 3.4.1 Wind analysis We used ASCE 7-05, method 2 to calculate the wind pressures. Table 2 lists the various parameters used. We considered the four cases depicted in figures 6–9 of ASCE 7-05. Each of these cases includes the direct wind force in one of four orthogonal directions either alone or in combination with storey torsional moment caused by the eccentric application of this load. We calculated the gust effect factor (Gf) following section 6.5.8.2 of ASCE 7-05 for flexible or dynamically sensitive structures with 1% assumed damping. We calculated wind base shears of 1436 and 2629 kip, respectively, along the two orthogonal directions. 3.4.2 Seismic analysis We used linear response spectrum analysis to calculate seismic forces and displacements. Per ASCE 7-05, Cl 12.9.4, we scaled the forces to 85% of the base shear obtained from the equivalent lateral force procedure of Cl 12.8. Table 3 lists the various parameters that we used for the seismic design. Table 1. Gravity loading criteria Description/Location Roof Mechanical, electrical at roof Residential including balconies Corridors, lobbies and stairs Retail Parking garage, ramp Construction loading Cladding Superimposed dead Live load Reducability 28 psf Total of 100 kip 28 psf 28 psf 110 psf 3 psf 3 psf 15 psf 25 psf – 40 psf 100 psf 100 psf 40 psf1 30 psf – Yes – Yes No No Yes No – PEER document showed 50 psf. SGH considered 40 psf in keeping with ASCE 7-05. Table 2. Wind design criteria Parameter Value Basic wind speed, 3 s gust (V) Basic wind speed, 3 s gust (V), for serviceability wind demands based on a 10-year mean recurrence interval Exposure Occupancy category Importance factor (Iw) Topographic factor (Kzt) Exposure classification Internal pressure coefficient (GCpi) Mean roof height (h) Wind base shear along two orthogonal directions Copyright © 2009 John Wiley & Sons, Ltd. 85 mph 67 mph B II 1·0 1·0 Enclosed ±0·18 544 ft, 6 in. 1436 and 2629 kip Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 81 ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB Table 3. Seismic design parameters Parameter Value Building latitude/longitude Occupancy category Importance factor (Ie) Spectral response coefficients Seismic design category Lateral system Undefined II 1·0 SDS = 1·145; SD1 = 0·52 D Buckling-restrained braced frames, non-moment-resisting beam column connections 7 5·5 2·0 3·16 s 0·051 W (governed by Cs-min from equation 12.8-5) 3504 kip (85% of static base shear) Modal response spectral analysis Response modification factor (R) Deflection amplification factor (Cd) System overstrength factor (Ω0) Building period (T) using Cl. 12.8.2 Seismic response coefficient Cs (Eq. 12.8-1) Scaled spectral base shear Analysis procedure Actual period from dynamic model: TY = 5·05 s; TX = 3·62 s. Spectral Acceleration, Sa (g) 1.6 1.2 0.8 0.4 0 0 2 4 6 8 Period (sec) Figure 2. Response spectrum for code analysis We performed the lateral analysis using ETABS, version 9.5.0. We used the 5% damped, acceleration response spectrum shown in Figure 2. We included 120 modes to obtain participation of at least 90% of the structure’s mass. We scaled the results of the response spectrum analysis such that the base shear immediately above the ground floor matched 85% of the static base shear from the equivalent static force analysis. Note that for levels below the ground floor, the mass of the perimeter walls is automatically calculated by the program. The 2007 California Building Code requires calculation of the redundancy factor (ρ) for seismic design. In accordance with ASCE 7-05 section 12.3.4.2, buildings qualify for a value of ρ = 1 provided that the removal of an individual brace or connection will not result in more than a 33% reduction in storey strength, nor will the resulting system have an extreme torsional irregularity. Extreme torsional irregularity is defined to exist when the maximum storey drift computed including accidental torsion, at one end of the structure transverse to an axis, is more than 1·4 times the average storey drifts at the two ends of the structure. Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 82 A. DUTTA AND R. O. HAMBURGER In order to calculate the redundancy factor for seismic forces acting in the north–south direction, we removed a brace from line 2 along the full height of the building. The code would actually permit this to be done one storey at a time; however, recognizing that we were being conservative, we removed the brace from all stories simultaneously, to reduce analysis effort. We then subjected the building to the static lateral forces along with the 5% accidental torsion. We term these forces as EQYPL (with the accidental eccentricity towards the positive right) and EQYMN (with the accidental eccentricity towards the negative left). Since the building has at least six bays of braced frame in each principal direction, loss of a brace or connection will never reduce the capacity by 33%. We found that for the north–south direction, the largest ratio of maximum storey drift over the average drift over the height of the building was 1·39 at storey 13. Since this is less than 1·4, we concluded that the redundancy factor in the north–south direction is 1. Figure 3(a) shows the elevation at line 2 after the removal of the brace. We performed a similar analysis in the X (east–west direction). In this analysis, we removed a single brace from line D and subjected the building to static forces EQXPL (with the accidental eccentricity towards the positive top) and EQXMN (with the accidental eccentricity towards the negative bottom). Figure 3(b) shows the elevation of the frame along line D with the brace removed. We then calculated the ratio of maximum to average drifts, and obtained ratios ranging from 1·468 to 1·408 from the roof to storey 36. However, section 12.3.4.2 stipulates that this ratio should strictly be studied for a floor resisting more than 33% of the base shear which is numerically equal to 1156 kip. This only happens below storey 35. In addition, as noted previously, our simultaneous removal of a brace at all levels is conservative. Since the ratio of maximum to average drift is only marginally greater than 1·4 at two stories with a maximum value of 1·408 at storey 34 (D/C of 1·008). We declared that this design met the requirements of ASCE 7-05 for a redundancy coefficient, ρ = 1. 3.5 Lateral force-resisting member design 3.5.1 Buckling-restrained brace (BRB) design We designed the BRBs for the worst of the wind and seismic loads. In almost all cases, seismic loads governed the design. Brace capacity in tension and compression was taken as φAsFy, with φ = 0·9 and Fy = 38 ksi. 3.5.2 Braced frame column design AISC’s Seismic Provisions for Structural Steel Buildings (ANSI/AISC 341-05) require columns in buckling-restrained braced frames to be checked for: (1) Axial load and moment interaction for code level forces (2) Axial load only corresponding to the sum of the vertical component of the strength of all BRBs that frame into the column along with tributary gravity loading We checked the columns for both of the above criteria, and found that the latter generally produced larger D/C ratios. Since the current configuration uses columns that form part of lateral framing in two orthogonal directions, we used the 100%–30% combination to calculate the maximum compression from the braces. We established the maximum compression forces from the brace as RyωβAsFy, where Ry = 1·1, ω = 1·25 and β = 1·1. We used built-up square box columns infilled with high strength ( f ′c = 10 ksi) concrete to resist the large compression force demands. The plate thicknesses of the box columns ranged from 1·5 in. at upper stories to 3 in. at lower ones. The plan dimension of the box columns ranged from 18 to 57 in. We calculated the strength, axial area and stiffness of infill columns using the provisions of chapter I Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 83 ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB (a) (b) Figure 3. Elevation of frame along lines 2 and D following removal of a brace. (a) Line 2 and (b) line D (‘Design of Composite Members’) of the 13th edition LRFD Steel Code (ANSI/AISC 360-05). We included this enhancement of axial area and stiffness over the bare steel section properties in the ETABS analysis model as well. We did this by defining a steel box section in ETABS and then using the property modifier to account for the infill concrete. Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 84 A. DUTTA AND R. O. HAMBURGER 3.5.3 Braced-frame beam design We designed the braced-frame beams for the horizontal component of the brace compression along with the unbalanced upward component of the BRB given by RyωAsFy(β-1)sinα, where α is the brace angle. Since BRBs are always stronger in compression than in tension, this moment is opposite in direction to the gravity moment. 3.5.4 Storey drifts for code design We calculated inter-storey drifts using response spectrum analysis. ASCE 7-05 stipulates the use of a response spectrum that is reduced by the response reduction factor R. It also states that the drifts obtained from this reduced spectral analysis be amplified by Cd to yield design storey drifts. Table 4 shows the design inter-storey drifts (amplified by Cd) obtained at various levels both for the X and Y directions. As can be seen, the maximum storey drifts are less than the permissible value of 0·02 at all levels. 4. PERFORMANCE-BASED DESIGN We based the seismic design for this alternate on the seismic design criteria (dated 29 April 2008) published by the Los Angeles Tall Buildings Structural Design Council (LATBSDC 2008). Since this design was not required to meet all of the prescriptive criteria contained in the building code, we were able to reduce the size and number of bays of the buckling-restrained frames. Specifically, along lines 2 and 7, we omitted two bays of bracing below level 10. Table 5 lists the performance objectives we followed for this design. Table 4. Storey drifts for code design Storey 40 39 38 37 36 35 34 33 32 31 30 29 28 27 26 25 24 23 22 21 20 19 Y drifts X drifts Storey Y drifts X drifts 0·013 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·014 0·013 0·013 0·013 0·013 0·012 0·012 0·012 0·012 0·011 0·005 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 0·006 18 17 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 0·011 0·011 0·010 0·010 0·010 0·009 0·008 0·007 0·004 0·004 0·004 0·004 0·004 0·004 0·003 0·003 0·003 0·002 0·006 0·006 0·006 0·006 0·005 0·005 0·005 0·005 0·004 0·004 0·004 0·004 0·004 0·004 0·004 0·004 0·003 0·002 Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 85 ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB Table 5. Seismic performance objectives Level of earthquake Earthquake performance objectives Frequent/Service: 25-year return period, 2·5% damping Maximum considered earthquake (MCE): As defined by ASCE 7-05, section 21.2, 2·5% damping Serviceability: Essentially elastic performance with minor yielding of brbf-s. Drift limited to 0·5% Collapse prevention: Extensive structural damage, repairs are required and may not be economically feasible. Drift limited to 3% Spectral Acceleration, Sa (g) 0.4 0.3 0.2 0.1 0 0 2 4 6 8 Period (sec) Figure 4. Response spectrum for serviceability analysis 4.1 Service level earthquake evaluation We used response spectrum analysis for this evaluation. We used the response spectrum shown in Figure 4, which was provided by PEER. The damping level was set at 2·5%. We neglected accidental torsion. Inter-storey drifts obtained for the service level earthquake are shown in Table 6. As can be seen, the drifts are less than 0·5% at all levels. We also evaluated the building for the same wind forces that we used for the code-based design described in the previous section. We found that wind always governed the required brace strength as compared with service level earthquake demands. 4.2 Maximum considered earthquake evaluation As required by the LATBSDC document, we performed a nonlinear response history analysis of the building for maximum considered earthquake shaking as defined by the building code. We used seven pairs of acceleration histories provided to us by PEER for this purpose. We used CSI Perform, version 4.0.1 with a constant modal damping level of 2·5% and 0·1% Rayleigh damping. We explicitly modelled nonlinearities in the BRBs. We modelled columns and beams as elastic elements, and later verified that demands on these elements remained within their elastic capacities. Details of the element modelling are discussed below. 4.2.1 Modelling of BRBs We modelled the BRBs using the perform BRB inelastic element in series with an elastic spring that models the end attachment of the braces. The BRB backbone curves were modelled as shown in Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 86 A. DUTTA AND R. O. HAMBURGER Table 6. Storey drifts for service level earthquake Storey Y drifts X drifts Storey Y drifts X drifts 40 39 38 37 36 35 34 33 32 31 30 29 28 27 26 25 24 23 22 21 20 19 0·0031 0·0032 0·0033 0·0033 0·0034 0·0033 0·0033 0·0033 0·0032 0·0031 0·0031 0·0029 0·0028 0·0027 0·0026 0·0026 0·0025 0·0024 0·0023 0·0022 0·0021 0·0021 0·0014 0·0015 0·0016 0·0017 0·0017 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0018 0·0017 0·0017 0·0017 18 17 16 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1 0·002 0·002 0·0019 0·0019 0·0018 0·0018 0·0017 0·0017 0·0016 0·0016 0·0016 0·0015 0·0015 0·0015 0·0014 0·0014 0·0012 0·0008 0·0017 0·0017 0·0017 0·0016 0·0016 0·0016 0·0016 0·0016 0·0017 0·0017 0·0017 0·0017 0·0017 0·0017 0·0017 0·0016 0·0014 0·001 1.25%Ko ωRyFyAs RyFyAs Tension 20Δy 10Δy Ko Δy Δy 10Δy 20Δy Compression RyFyAs βωRyFyAs Figure 5. Presumed backbone curve for buckling-restrained braces Figure 5 with Ry = 1·1, ω = 1·25 and β = 1·1. Figure 6 shows a screen shot of the Perform input form for the BRB definition. Note that the initial stiffness (K0) of the BRB is based on AsE/L with L equal to 70% of the actual centre-to-centre length of the brace. The remaining 30% of the length is modelled as an essentially rigid element. 4.2.2 Modelling of the columns and beams We modelled the columns using the non prismatic steel sections in Perform. The moment of inertia input in the form was adjusted to account for the additional stiffening due to the presence of the infill concrete. The axial load and moment strength for the various columns were defined in Perform using Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB 87 R yF yA s 70% Actual 1.25% Ko Figure 6. BRBF property definition in Perform Figure 7. Column strength definition in Perform Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 88 A. DUTTA AND R. O. HAMBURGER the same process as for the elastic analysis. We assumed nominal strengths for the columns without any strength reduction or φ factors. Figure 7 shows the column axial load a representative flexural strength definition form in Perform. The compound column element associated with the geometric location of the column utilizes the cross-sectional information and assembles the column element as an elastic element. We have ensured that the column element always stays elastic during the analysis by monitoring the axial flexure interaction ratios so that they are always less than 1. Beams are modelled in an identical fashion as the columns, except that standard W sections are used. Similar to the columns, the beams are also modelled as elastic elements. 4.2.3 Modelling of the diaphragms at upper, ground and basement levels We modelled rigid diaphragms at all elevated floors. However, we modelled diaphragms at the ground and all basement floors with elastic shell elements using 30% of the gross cross-section properties. We considered an effective thickness comprising the total thickness of the topping plus half the rib thickness. Thus, the ground floor was modelled as 10·5-in. thick slab as it represents a 9-in. topping and 1·5-in. half rib height. We followed the same approach for the basement slabs except that the modulus of elasticity also included the λ modifier corresponding to lightweight concrete. Peak Interstory Drift in Y Direction 40 35 30 Story 25 20 15 10 δy 5 0 -5 0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 Peak Interstory Drift (%) Average Denali Ps Landers Yermo LP Gilroy LP Saratoga Nridge Sylmar CS Nridge Sylmar Hosp Parachute Figure 8. Peak inter-storey drifts in the Y direction Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 89 ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB 4.2.4 Modelling of perimeter walls We modelled the perimeter shear walls as elastic wall elements with cracked stiffness equalling 50% of the gross stiffness. The elastic shear modulus was considered to be 40% of the gross elastic (Young’s) modulus per ASCE 41, supplement 1. 4.2.5 Selected results from nonlinear analysis 4.2.5.1 Storey drifts Figures 8 and 9 present the storey drift profiled obtained from the analyses. Figure 8 shows the drift in the Y direction and Figure 9 the X direction. The maximum drift in the Y direction is around 1·97%, while the maximum in the X direction is at 2·25%. As can be observed from the figures, the drifts are less than the permissible value of 3% in both directions, and the mean peak drifts are generally on the order of 1·25% or less. 4.2.5.2 Column axial flexure interaction ratio Since the beam columns were modelled as elastic elements, we monitored the interaction ratios closely and ensured that they remain less than unity signifying elastic behaviour. We found that the maximum mean interaction ratio was 0·63. The absolute maximum from the seven analyses is 0·83, justifying our assumption regarding the elastic behaviour of the columns. Peak Interstory Drift in X Direction 40 35 30 Story 25 20 15 10 5 δx 0 -5 0.00 0.50 1.00 1.50 2.00 2.50 Peak Interstory Drift (%) Average Denali Ps Landers Yermo LP Gilroy LP Saratoga Nridge Sylmar CS Nridge Sylmar Hosp Parachute Figure 9. Peak inter-storey drifts in the X direction Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 90 A. DUTTA AND R. O. HAMBURGER Line 2 Line 3 Line 4 Line D Figure 10. BRBF core strain DCRs along lines 2, 3, 4 and D Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 91 ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB Accumulated Peak Story Shear in Steel Superstructure in Y Direction 40 Vy 35 30 Story 25 20 15 10 5 0 -5 0 2500 5000 7500 10000 12500 Accumulated Peak Story Shear (kips) Average Denali Ps Landers Yermo LP Gilroy LP Saratoga Nridge Sylmar CS Nridge Sylmar Hosp Parachute Figure 11. Accumulated peak storey shear in steel superstructure in Y direction 4.2.5.3 BRB core strain We also monitored strain in the BRB cores, and ensured that the mean strain is less than of 0·013 (10εy) obtained from the test results conducted at the University of Utah by Romero and Reavely. Figure 10 shows the BRB core strain for lines 2, 3, 4 and D, respectively. The results are presented as D/C ratio with the capacity being 0·013. The ratios are all less than unity, and most are substantially less than unity, implying acceptable performance. DCRs indicated in yellow in the figure indicate values in excess of 75% of the 0·013 strain limit. 4.2.6 Storey shears We also monitored storey shears over the building height. Figures 11 and 12 plot the storey shears in the Y direction and X direction, respectively. The effect of the stiff ground floor diaphragms is easily observable from the plots. We have proportioned the ground floor slab and the collectors so that it is capable of transferring this large reaction. 5. SUMMARY AND CONCLUSIONS Simpson Gumpertz & Heger performed two alternative designs of a 40-storey building located at a hypothetical site in Los Angeles. Both buildings utilize buckling-retrained braced frames as the lateral Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal 92 A. DUTTA AND R. O. HAMBURGER Accumulated Peak Story Shear in Steel Superstructure in X Direction 40 35 Vx 30 Story 25 20 15 10 5 0 -5 0 2000 4000 6000 8000 10000 Accumulated Peak Story Shear (kips) Average Denali Ps Landers Yermo LP Gilroy LP Saratoga Nridge Sylmar CS Nridge Sylmar Hosp Parachute Figure 12. Accumulated peak storey shear in steel superstructure in X direction system. The first design strictly follows the provisions of the 2007 California Building Code except that it exceeds code limitations on the height for structures not having backup special moment frames. The columns and beams are designed using the relevant requirements of AISC 341.05 seismic provisions. The second design uses the concepts of performance-based engineering and is based on the LATBSDC. A nonlinear time history analysis is performed on this design using maximum credible earthquake input to ensure a safe design. Generally, the second design produced fewer bays of bracing, smaller sizes of the BRBs and columns. The main savings, however, is likely to be from the design of the foundation. The code design requires that the foundation be designed for the accumulated vertical component of the brace forces over the height of the building. This essentially implies exclusively first-mode behaviour which does not occur in tall buildings. Using the actual demands from the time history analysis, we obtained forces that are smaller and hence more manageable than the code-required forces. REFERENCES (American Institute of Steel Construction). ANSI/AISC 341-05, 2005 Seismic Provisions for Structural Steel Buildings. AISC: American Institute of Steel Construction, Chicago, IL. Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal ARCHETYPICAL DESIGNS FOR A 40-STOREY BRB 93 ANSI/AISC 360-05, 2005, Specification for Structural Steel Buildings, American Institute of Steel Construction, Chicago, IL. ASCE 2005. Minimum Design Loads for Buildings and Other Structures, American Society of Civil Engineers, Reston, VA. LATBSDC (Los Angeles Tall Buildings Structural Design Council). 2008. An Alternative Procedure for Seismic Analysis and Design of Tall Buildings Located in the Los Angeles Region. LATBSDC: Los Angeles, CA. Copyright © 2009 John Wiley & Sons, Ltd. Struct. Design Tall Spec. Build. 19, 77–93 (2010) DOI: 10.1002/tal