Seismic design and testing of buckling‐restrained braces with a thin profile

Abstract

Summary A thin‐profile buckling‐restrained brace (thin‐BRB) consists of a rectangular steel casing and a flat steel core that is parallel to a gusset plate. A thin configuration reduces the width of the restraining member and thus saves usable space in buildings. However, deformable debonding layers, which cover the steel core plate in order to mitigate the difference between the peak tensile and compressive axial forces, provide a space for the steel core to form high mode buckling waves when the thin‐BRB is under compression. The wave crests squeeze the debonding layers and produce outward forces on the inner surface of the restraining member. If the restraining member is too weak in sustaining the outward forces, local bulging failure occurs and the thin‐BRB loses its compression capacity immediately. In order to investigate local bulging behavior, a total of 22 thin‐BRB specimens with a ratio of steel core plate to restraining steel tube depth ranging from 0.3 to 0.7 and axial yield force capacities ranging from 421 kN to 3036 kN were tested by applying either cyclically increasing, decreasing, or constant axial strains. The restraining steel tube widths of all the specimens were smaller than 200 mm and were infilled with mortar with a compressive strength of 97 MPa or 55 MPa. Thirteen of the 22 thin‐BRB specimens' restraining members bulged out when the compressive core strains exceeded 0.03. A seismic design method of the thin‐BRB in preventing local bulging failure is proposed in this study. Test and finite element model (FEM) analysis results suggest that the outward forces can be estimated according to the BRB compressive strength, steel core high mode buckling wavelength, and the debonding layer thickness. In addition, the capacity of the restraining member in resisting the outward forces can be estimated by using the upper bound theory in plastic analysis. Both the FEM analysis and test results indicate that the proposed method is effective in predicting the possibility of local bulging failure. Test results indicate that the proposed design method is conservative for thin‐BRB specimens with a large steel core plate to restraining steel tube depth ratio. This paper concludes with design recommendations for thin‐BRBs for severe seismic services. Copyright © 2015 John Wiley & Sons, Ltd.

Seismic performance analysis of BRBs and gussets in a full‐scale 2‐story BRB‐RCF specimen

“…This study considers the effects of the infilled grout on the restrainer's axial and rotational stiffness values: where α is a factor for reducing the elastic stiffness of the infilled grout. Based on previous BRB tests, severe cracking of the infilled grout was observed due to the high‐mode buckling of the BRB core. Thus, it appears appropriate to reduce the grout's contribution.…”

“…In addition, the shell elements are very effective in simulating the in‐plane or OOP force vs deformation characteristics of steel plates . The BRB core high mode buckling has been studied by using solid elements for the steel core . In these studies, the effects of frictions developed between the BRB core and the restrainer including steel casing and infill mortar were incorporated.…”

“…24 The BRB core high mode buckling has been studied by using solid elements for the steel core. 25,26 In these studies, the effects of frictions developed between the BRB core and the restrainer including steel casing and infill mortar were incorporated. Numerical instability could be encountered in simulating the hard contact and friction characteristics.…”

Energy-Based Design Criterion of Dissipative Bracing Systems for the Seismic Retrofit of Frame Structures

“…The elastic limit values of the level shear F e (i.e., the sum of the elastic limit shear forces of the columns) and the first interlevel drift ILD e (replacing ID e in this case) computed in X direction, named F e,X and ILD e,1L,X , are equal to 969 kN and 22 mm, respectively. Introducing these values, as well as α F and ξ eq,αF values given by (25) and (26), in (18), the following E D estimate is derived: E D,αF = 2πα F F e,X ξ eq,αF ID e,1L,X = 65.6 kJ (27) Dividing E D by the number of spring-dampers placed in X, the minimum energy dissipation capacity E D,X,d to be assigned to each of the eight devices in order to reach the target performance at the MCE results as follows: E D,X,d = 8.2 kJ. The spring-damper type with the nearest nominal energy dissipation capacity, E n , to E D,X,d has the following mechanical properties, drawn from the manufacturer's catalogue [48]: E n = 9 kJ; stroke s max = ±30 mm; damping coefficient c = 9.9 kN(s/mm) γ , with γ = 0.15; F 0 = 17 kN; and k 2 = 1.74 kN/mm.…”

“…At the same time, new technologies, or improved versions of existing ones, have been implemented in the field of energy dissipation, capable of supplying supplemental damping and horizontal stiffness in different proportions, depending on the mechanical characteristics of dampers and their installation layout. Among the latest achievements in this field, systems incorporating updated models of metallic yielding devices like ADAS (Added Damping and Stiffness) and TADAS (Triangular Added Damping and Stiffness) elements [8][9][10][11][12], shear panels [13][14][15][16][17][18], shear panels with openings [19][20][21][22][23], or buckling-restrained braces [24][25][26][27], have been proposed, which typically provide significant contributions in terms of both properties. Viscous dampers can offer stiffening effects too, in addition to supplemental damping [28][29][30].…”