Ductile Fuse for Special Concentrically Braced Frames and Related Methods

Abstract

Embodiments of the present disclosure provide a structural improvement in braced frame structures. An exemplary solution involves heat-treating a mid-section region of a steel brace, the steel brace comprising hollow structural sections tubing; after heat-treating the mid-section region, cooling the mid-section region of the steel brace; wherein the heat-treated steel brace has changed mechanical properties due to the heat-treatment, the changed mechanical properties including improved material ductility, work hardening ability, and notch toughness; and coupling the heat-treated steel brace to a gusset plate within a braced building frame structure without need of gusset plate reinforcement.

Description

BACKGROUND

Due to their inherently high lateral stiffness, special concentrically braced frames (SCBFs) generally lead to economical building frames. Over the past two decades several advances have been made in understanding the seismic performance and failure modes of SCBFs. These efforts have led to among others, improvements in connection detailing, restrictions on brace configuration and more stringent brace slenderness and section compactness criteria. However, despite these advancements, areas for improvement in the economy and performance of SCBFs still exist.

BRIEF DESCRIPTION OF THE DRAWINGS

Many aspects of the present disclosure can be better understood with reference to the following drawings. The components in the drawings are not necessarily to scale, emphasis instead being placed upon clearly illustrating the principles of the present disclosure. Moreover, in the drawings, like reference numerals designate corresponding parts throughout the several views.

FIG. 1a is an image showing a rupture of a brace to a gusset plate connection.

FIG. 1b is a diagram of a sketch of a reinforced connection between a brace and a gusset plate.

FIG. 1c is a diagram of a sketch of a heat-treated braced frame (e.g., an inverted V-braced frame with heat-treated braces) in accordance with various embodiments of the present disclosure.

FIGS. 2a-2b are diagrams showing tensile stress-strain responses of A500 steel and heat-treated (HT) A500 steel at various peak temperatures and cooling temperatures in accordance with embodiments of the present disclosure.

FIGS. 2c-2d show optical micrographs of A500 and heat-treated A500 steel in accordance with embodiments of the present disclosure.

FIGS. 2e-2f are images of saw cuttings on a hollow structural section before and after heat treatments in accordance with embodiments of the present disclosure.

FIG. 2g is an image of a fractured Charpy V-notch bars before and after heat treatments in accordance with embodiments of the present disclosure.

FIG. 3a shows a table for a chemical composition of A500 steel in accordance with the present disclosure.

FIG. 3b shows a table for mechanical properties of A500 and HT A500 steels in accordance with the present disclosure.

FIG. 3c shows a table of results of Finite Elements analysis for A500 and HT A500 steels in accordance with the present disclosure.

FIG. 4a is a diagram showing a far field loading history used in testing a heat-treated A500 steel in accordance with embodiments of the present disclosure.

FIG. 4b is a diagram of an embodiment of a field element (FE) mesh and boundary conditions for an exemplary heat-treated brace in accordance with the present disclosure.

FIG. 4c is a diagram showing images of cyclic and smooth notch tensile (SNT) tests along with a contour plot of hydrostatic stress distribution for heat-treated A500 steel in accordance with embodiments of the present disclosure.

FIG. 4d is a diagram showing experimental and simulated stress-strain response of A500 and heat-treated A500 steels in accordance with embodiments of the present disclosure.

FIG. 4e is a diagram of experimental and simulated SNT responses of A500 and heat-treated A500 steels in accordance with embodiments of the present disclosure.

FIGS. 5a-5b are diagrams showing a comparison of experiment and simulated force-drift responses of a square hollow structural section (HSS) tubing.

FIG. 5c is a cyclic void growth model (CVGM) fracture prediction for an HSS tubing.

FIG. 5d is a field element (FE) prediction of accumulated equivalent plastic contours of an HSS tubing in accordance with embodiments of the present disclosure.

FIG. 5e is an image showing a fracture initiation of an HSS tubing.

FIG. 6a is a diagram showing a force-drift response of an HSS tubing that has not been heat-treated in accordance with an exemplary technique of the present disclosure.

FIG. 6b is a diagram showing a force-drift response of an HSS tubing that has been heat-treated at 700° C. in accordance with an exemplary technique of the present disclosure.

FIGS. 6c-6d are diagrams showing equivalent plastic strain contours of an HSS tubing (A500) at 2.7% drift before and after heat-treatment at 700° C. in accordance with an exemplary technique of the present disclosure.

FIGS. 6e-6f are diagrams showing equivalent plastic strain contours of an HSS tubing (A500) at 4% drift before and after heat-treatment at 700° C. in accordance with an exemplary technique of the present disclosure.

DETAILED DESCRIPTION

In special concentrically braced frames, brace yielding and buckling are the primary mechanisms of seismic energy dissipation, and mid-length brace fracture is the preferred failure mode. However, as evidenced by several studies, the connections between the braces and gusset plates are susceptible to premature fracture as are other connections e.g. between the gusset plate and beam or column. The present disclosure describes the development of a ductile fuse to prevent such failures. Embodiments of such a fuse are created by selectively reducing the brace material strength through a thermal treatment process which in turn reduces connection force demands. Material testing of ASTM A500 steel from square hollow structural sections (HSS) tubing was conducted to characterize the mechanical property changes from the thermal treatment and to calibrate 3D finite element (FE) models of brace members. The experimental data and FE models demonstrate the benefits of the fuse, which include reduced tensile force demands at the brace to frame connections, improved fracture resistance of brace material, increased drift capacity, residual stress relief, and improved element compactness.

Rectangular or round hollow structural sections (HSS) constructed to the ASTM A500 grade B or C specification are commonly specified for brace members due to their structural efficiency, availability, and ease of construction. However, since braces are sized using design strengths based on the specified minimum yield strength and connections, and other elements are designed to resist forces based on the expected yield strength, the high ratio of expected to specified yield strength (R_y), which is 1.3-1.4 for A500 steel, compromises economy. Stated more directly, connections and other frame elements such as columns are designed to resist 30 to 40% higher brace forces than otherwise necessary because of brace material over strength.

Another issue is that cold forming at the corners of rectangular HSS members reduces material ductility and notch toughness. Unfortunately, this is also the location of high strain demands when mid-length brace hinges are formed under large load reversals. These combined effects likely reduce the deformation capacity of rectangular HSS braces under seismic loading.

Additionally, HSS braces are typically connected to surrounding elements (column, beam, or both) with a gusset plate. The construction of this connection involves slotting the brace, thereby creating a net section susceptible to premature failure (as shown in FIG. 1a). Typically, the net section strength is determined as

R _{net section}=σ_uUA_n  (1)

where σ_u is the ultimate stress, U is the shear lag factor, and A_n is the net section area. While the yield strength of the member is

R _yield =R _yσ_yA_g  (2)

where σ_y is the yield stress and A_g is the section area. Typically, the shear lag factor, U, is approximately 0.85, and the ratio between the yield and the ultimate stress is also a similar value. Thus, even in absence of material over strength (i.e. R_y=1) the inequality σ_uUA_n>R_yσ_yA_g, cannot be satisfied unless the net section is reinforced with plates as shown in FIG. 1b.

To overcome these challenges, a technique is deployed. As shown in FIG. 1c, an exemplary technique of the present disclosure involves heat-treating the mid-section of the brace (up to one brace depth away from the edge of gusset plate) by subjecting this region to high temperatures followed by controlled cooling. Temperatures during this heat-treatment are high enough to relieve residual stresses from prior cold work and to cause metallurgical changes to the material. Additional details on elements of the exemplary heat-treatment technique as applied to moment resisting frames is available in an article by Morrison, M., Schweizer, D., and Hassan, T. (2015), entitled “An innovative seismic performance enhancement technique for steel building moment resisting connections,” J. Const. Steel Res., 109(6), 34-46, which is incorporated herein by reference in its entirety.

This technique has several attractive features for braced frames. In FIG. 2a and FIG. 2b the tensile stress-strain responses of A500 steel and heat-treated (HT) A500 steel are shown. Coupons were subjected to various peak temperatures (700, 800, and 900° C.), held at this temperature for 30 minutes (to allow for thermal equilibrium), cooled at 6° C./min (FIG. 2a) or 0.6° C./min (FIG. 2b) to 550° C., and then removed from the furnace to cool in air. From the figures, it is observed that with increasing temperatures and slower cooling rates the strength of steel is reduced while the elongation is increased. Also note that the ability of the steel to work harden is restored. Temperature has a far greater effect than cooling rate, as such, for the remainder of this disclosure, discussion will be limited to the thermal history shown in FIG. 2a (i.e. the faster cooling rate).

The reduction in strength and increase in ductility are due to microstructural transformations (recovery and partial or full recrystallization) which take place during the annealing or heat-treating cycle. A brief overview of these metallurgical processes is provided here. Recrystallization of a cold worked or rolled metal reduces the dislocation density in the crystal lattice and produces a new set of “strain-free” grains. The process utilizes mass transport (i.e. diffusion), making it especially sensitive to temperature (typically following an Arrhenius type equation). Below the phase transformation temperature e.g. at 700° C. (Ac1≈717° C. for this steel), the thermodynamic driving force for recrystallization is the free energy stored in dislocation networks and the process allows the material to achieve a lower internal energy state by reducing the strain energy in the lattice. At higher temperatures, recrystallization is due to the ferrite to austenite phase transformation, which reverses upon cooling. Because of recrystallization, the effects of prior cold work are removed either partially or completely. After full-recrystallization (e.g. HT at 900° C.) a uniform equiaxed grain structure is produced (compare FIG. 2c and FIG. 2d). Note that there is no significant grain growth, this is likely due to the solute drag and Zener pining effects from alloying elements such as Niobium and Aluminum (see Table 1 (FIG. 3a) for the chemical composition of this steel).

As a consequence of these microstructural changes, notch toughness is remarkably improved and residual stresses in the HSS section are relieved (see FIGS. 2e-2g and Table 2 (FIG. 3b)).

These mechanical property changes have several desirable effects for braced frames. First and foremost, material overstrength can be reduced significantly, for example as shown in FIG. 2a and Table 2 (FIG. 3b), heat-treating at 800° C. reduces the yield strength to 345 MPa which is the specified minimum yield strength of the material. Further, material strength can be controlled to allow the designer more flexibility in proportioning members. This will improve structural efficiency and likely overall economy of the structure. Second, the combination of improved ductility, notch toughness, and work hardening ability should reduce the susceptibility of braces to strain localization and fracture, allowing them to sustain larger deformations, more cycles and, better energy dissipation. Third, reinforcement of the connection between the brace and gusset plate (see FIG. 1a) is no longer needed since the inequality σ_uUA_n>R_yσ_yA_g can be readily satisfied for HT braces. Finally, residual stress relief should lead to improved compressive strength.

It is noted that recently the ASTM A1085 specification for cold rolled or formed HSS has been introduced and has been given a lower overstrength factor (R_y=1.25) than A500 steel. While improved performance is anticipated for the A1085 grade, the heat-treatment described in this disclosure can further reduce material overstrength and provide better ductility and notch toughness than A1085. In addition, it solves the problem of connection reinforcement, which is not addressed by the A1085 specification.

In the following discussion, the fuse technique is analyzed through detailed finite element (FE) simulations and an uncoupled ductile fracture prediction model by focusing on the effect of tensile properties and width/thickness (b/t) ratio on brace strength and deformation limit states such as the onset of local buckling and ductile fracture initiation. Details of the modelling strategy and results are presented in the following.

Fe Modelling

The modelling strategy employed in this disclosure follows closely with that of Fell, B.V. (in “Large-Scale Testing and Simulation of Earthquake-Induced Ultra Low Cycle Fatigue in Bracing Members Subjected to Cyclic Inelastic Buckling,” Ph.D. Dissertation, University of California-Davis, 2008) with a few exceptions. Three-dimensional nonlinear FE models were developed for brace members using the commercial software ANSYS Mechanical ADPL. Geometric and material nonlinearities were incorporated in the FE models. An example of the FE mesh and boundary conditions is shown in FIG. 4b. The brace and gusset plates were modeled with continuum 20 noded solid hexahedral elements (SOLID186). FE models accounted for material nonlinearity through rate-independent metal plasticity theory based on the von Mises yield criterion, additive strain decomposition, and associated flow rule. A multilinear kinematic hardening model calibrated from tensile material tests, strain-controlled cyclic material tests and smooth notch tensile (SNT) tests was used to capture material strain hardening (FIG. 2a and FIGS. 4a-4e). The strain-controlled cyclic material tests allowed for model calibration at low strain ranges while the tensile and SNT tests allowed for simulation of behavior at large strains. This combination was essential for capturing the response of the brace throughout the load history (up to fracture initiation).

Geometric nonlinearities were accounted for via a large deformation formulation which accompanied by small eccentricities/imperfections in the geometry allowed for the simulation of global and local brace buckling. Initial imperfections were obtained by first conducting an eigenvalue buckling analysis of the perfect structure and then prescribing superimposed scaled values of the global buckling mode (first eigenmode) and local buckling mode (third eigenmode) displacement fields as the initial configuration of the structure. The maximum value of geometric imperfection for the global mode was L/1000, which is half of the permissible variation in straightness allowed by ASTM A500. The maximum value of geometric imperfection for the local mode was 0.07 mm based on measurements by Fell (2008). Global and local imperfections were essential to simulating buckling deformations observed in experimental testing of braces. It is noted, that residual stresses were not included in the FE analysis leaving the effect of this variable available for further analysis.

Stress and strain indices from the FE model were used as inputs to a micromechanics based cyclic void growth model (CVGM) proposed by A. M. Kanvinde and G. G. Deierlein (in “Cyclic void growth model to assess ductile fracture initiation in structural steels due to ultra-low cycle fatigue.” Journal of engineering mechanics 133(6), 701-712, 2007) to predict the initiation of ductile macroscopic cracks that were reported during laboratory tests by Fell (2008). The CVGM predicts the initiation of a ductile macroscopic crack when the following condition is satisfied

VGI _cyclic >VGI _{cyclic critical} f or l≥l*  (3)

VGI _cyclic=Σ_{Tensile Cycles}∫e^1.5Tdε_P−Σ_{Compressive Cycles}∫e^1.5Tdε_P  (4)

VGI _{cyclic critical} =VGI _{critical monotonic} −e ^{λε p accumulated}   (5)

where, VGI_{critical monotonic} is obtained from the following failure criterion establish for monotonic loading:

VGI _monotonic=∫₀^εpe^1.5Tdε_p>VGI_{critical monotonic}  (6)

where T=σ_m/σ_e (ratio of mean stress to effective stress, also called stress triaxiality) and dε_p is the increment of equivalent plastic strain. λ is a material dependent parameter which represents the degradation of material resistance to ductile fracture due to cyclic loads and ε_p^accumulated is the equivalent plastic strain that has accumulated up to the beginning of each tensile excursion of loading. The CVGM fracture initiation criterion (Eq. 3) should be satisfied over a length scale representative of the physical events leading to ductile fracture (l*). This model feature typically employs mesh sizes on the order of l*. The element length in the region of interest is 0.3 cm which is sufficient to capture strain and stress gradients but is much larger than l*. However, other studies demonstrate reasonable results using similar element sizes.

VGI_{critical monotonic} and λ are calibrated by experimental testing of notched bar specimens subjected to monotonic and cyclic loading histories respectively. In this analysis, SNT bars subjected to monotonic loading were used to calibrate VGI_{critical monotonic} (see FIG. 4e and Table 2 (FIG. 3b)) but due to limited resources, cyclic notched bar testing was not carried out. Therefore, a value of λ=0.17 was used based on the work of Fell (2008). While this value was found to be reasonably accurate for A500 steel, it may be conservative for HT A500 steel.

Finite Element Model Validation

For FE model validation, specimen HSS 1-1 and HSS 1-2 tested by Fell (2008) were modelled. Both specimens consisted of square HSS102×102×6.35 members. The overall length of the specimens (i.e. the distance between the outer edges of the gusset plates) was 3124.2 mm. The gusset plates were welded to thick endplates and loaded axially. Specimen HSS 1-1 was subjected to a far-field loading history (see FIG. 4a) akin to that commonly used in the testing of moment frames, while specimen HSS1-2 was subjected to a near-fault loading history with asymmetric compression loading cycles. Story drift angle expressed in radians (θ) is determine from the assumed kinematics of the braced frame during seismic loading and is calculated as

θ=2Δ_a/L_B  (7)

where Δ_a is the axial deformation of the brace and LB is the distance between the fold lines of the gusset plate (2984.5 mm).

The predicted brace force-drift (%) response is plotted against the experimental response in FIG. 5a and FIG. 5b. The model prediction is found to be in close agreement with the experimental response for both loading histories. The evolution of the void growth indices (for the CVGM) for specimen HSS 1-1 is shown in FIG. 5c. The location considered is the brace centerline at the corner of the HSS as shown in FIG. 5d. Note that fracture initiation is predicted to occur at the end of the second excursion of loading at 2.8% drift while fracture was observed slightly before the end of the second excursion at 2.8% drift (FIG. 5a).

Fe Evaluation of Exemplary Technique

An exemplary fuse technique of the present disclosure is evaluated using the previously validated FE model. See Table 3 (FIG. 3c) for details of the specimens considered.

The length of the brace and boundary conditions are the same as those used for FE model validation. Also, for sake of simplicity, the far field load history used in the validation is applied to all specimens. This evaluation considers the effect of the maximum temperature of heat treatment (i.e. the change in mechanical properties from heat-treating the brace to 700, 800, and 900° C. respectively). In addition, the evaluation considers two different HSS wall thicknesses to evaluate the effect of section compactness i.e. b/t ratio on brace response. Note that measured properties (from Table 2 (FIG. 3b)) are used for calculation of all applicable quantities reported in Table 3 (FIG. 3c). The yield strength from heat-treating at 700, 800, and 900° C. resulted in three different ratios of σ_uUA_n/σ_yA_g (capacity to demand) all of which are greater than unity. This implies that net section failure would be circumvented for all HT braces considered. The prediction of fracture initiation at the net section was not included in the FE analysis, but is available for further study.

Several important insights are drawn from the results of FE analysis presented in Table 3 (FIG. 3c). The normalized tensile (P_Tmax/P_y) and compressive strengths (P_Cmax P_y) of the brace are increased from the heat treatment. The increase in normalized compressive strength is far greater than the increase in tensile strength. The mild increase (3-9%) in normalized tensile strength when HT braces are compared to their unconditioned counterparts is likely due to material strain hardening (see FIG. 2a and Table 2 (FIG. 3b)).

The b/t ratios for the HSS members have been normalized to the limiting values specified in the AISC Seismic Provisions (at “Seismic provisions for structural steel buildings,” American Institute of Steel Construction, Chicago, USA, 2016) and tabulated in Table 3 (FIG. 3c). Note that the heat treatment reduces the normalized b/t ratios implying improved compactness and greater deformation capacity. Indeed, as shown in Table 3 (FIG. 3c), HT braces showed delayed onset of local buckling and fracture initiation. In fact, for HSS102×102×9.5 braces HT at 700, 800, and 900° C. no local buckling or fracture initiation was predicted before the analysis was terminated after loading up to 6% drift.

The improved deformation capacity of HT braces is attributed to the improved work hardening behavior and the increased fracture resistance of the material. For example, consider the responses of HSS102×102×9.5-A500 and HSS102×102×9.5-HT at 700° C. braces shown in FIGS. 6a-f. Due to the low work hardening rate of A500 steel, plastic strains localize in the center of the brace during cycles at 2.7% drift. Note the relatively large strain gradients at the center of the A500 brace shown in FIG. 6c relative to those of the HT brace shown in FIG. 6d. This localization of plastic strains eventually leads to local buckling during cycles at 4% drift and fracture soon after (FIG. 6a and FIG. 6e). This localization of strain is circumvented in the HT brace due to the improved material work hardening (FIG. 6b and FIG. 6f).

One exemplary heat-treatment of the present disclosure may employ a special furnace or induction heating system adjustable to the desired HT length of the brace. Note that heat-treatment of the brace can be performed before the brace is cut to its final length and before fabrication steps such as hole or slot cutting. Therefore, it will not interrupt the fabrication process. Also, based on the findings of this disclosure, the process may take approximately two hours, in various embodiments. This process time can be reduced with faster heat-up rates and further efficiency can be gained by heat-treating several braces simultaneously. Non-destructive test methods such as hardness testing can be calibrated to verify desired strength is achieved by the exemplary heat-treatment or alternatively, small pieces of steel from the same heat as the brace members can be HT and destructively tested to ensure strength requirements are satisfied.

While the exemplary fuse technique is shown to be novel during small scale testing, it may be further validated through large scale testing. Also, given the many variables that influence HSS material strength such as manufacturing history (e.g. direct formed vs. continuously formed), section geometry and steel chemistry, a detailed metallurgical study can be performed to develop appropriate thermal cycles that efficiently produce the desired mechanical properties. Fortunately, the kinetics and thermodynamics of recrystallization and phase transformations in low carbon steels are fairly well established. So too are metallurgically based strength prediction models. For example, based on the chemical composition (Table 1 (FIG. 3a)) and an average grain size of 11.7 μm for A500 steel HT at 900° C. (measured from FIG. 2d), one such strength model (by Totten G. E., Xie L., Funatani K., “Handbook of Mechanical Alloy Design” Marcel Dekker Inc., New York, 2004) predicts a yield strength of 279 MPa which is remarkably close to the average measured yield strength of 276 MPa (Table 2 (FIG. 3b)).

A novel fuse technique for SCBFs is introduced and analyzed in this disclosure. The technique involves heat-treating the mid-section of the brace to reduce material strength and improve work hardening, ductility and notch toughness. The technique has the potential to simultaneously reduce construction cost and improve seismic performance for new building construction and may also be used to upgrade older non-ductile brace frames.

Embodiments of the present disclosure provide a structural improvement in braced frame structures. An exemplary solution in accordance with the present disclosure deals with annealing (heat-treating) central portions of braces in steel concentrically braced building frames (CBBFs) by subjecting this region to high temperatures followed by controlled cooling. The technique reduces the yield and tensile strength, while increasing the ductility, work hardening ability, and notch toughness of cold rolled steel in hollow structural sections (HSS). These changes in mechanical properties have several desirable effects for CBBFs. First and foremost, material overstrength can be lowered. Consequently, the size of capacity protected elements such as beams, columns, and connections can be reduced making CBBFs more economical. Second, the combination of improved material ductility, notch toughness, and work hardening ability reduces the susceptibility of braces to strain localization and fracture, allowing them to sustain larger deformations prior to rupture. Third, reinforcement of the slotted connection between the HSS brace and gusset plate as currently required by AISC 341 is no longer needed since this connection is capacity protected by the weakened brace. Finally, residual stress relief from the annealing process should lead to improved brace compressive strength.

Therefore, such techniques in accordance with the present disclosure solve the problem of brace material overstrength in seismic design and construction of CBBFs; improve the ductility of braces for better seismic performance of CBBFs; and eliminate the need for brace to gusset plate reinforcement in CBBFs. Further, such techniques have immediate use in both new construction of CBBFs and retrofit of existing “non-ductile” CBBFs.

It should be emphasized that the above-described embodiments of the present disclosure are merely possible examples of implementations, merely set forth for a clear understanding of the principles of the disclosure. Many variations and modifications may be made to the above-described embodiment(s) without departing substantially from the spirit and principles of the present disclosure. All such modifications and variations are intended to be included herein within the scope of this disclosure and protected by the following claims.

Claims

1. A method comprising: heat-treating a mid-section region of a steel brace, the steel brace comprising hollow structural sections tubing;after heat-treating the mid-section region, cooling the mid-section region of the steel brace causing the heat-treated steel brace to change mechanical properties due to the heat-treatment, the changed mechanical properties including improved material ductility, work hardening ability, and notch toughness; andcoupling the heat-treated steel brace to a gusset plate within a braced building frame structure without need of gusset plate reinforcement.

2. The method of claim 1, further comprising coupling the gusset plate to a column or brace within the braced building frame structure.

3. The method of claim 1, wherein the steel brace comprises cold rolled steel.

4. The method of claim 5, wherein the steel brace is heat-treated at a temperature that relieves residual stresses from prior cold work on the steel brace.

5. The method of claim 1, wherein the mid-section region is heat-treated at a temperature of at least 700° C.

6. The method of claim 1, wherein the mid-section region is heat-treated at a temperature of 700° C. for 30 minutes.

7. The method of claim 6, wherein the mid-section region is cooled at 6° C./min before removing the steel brace from a furnace used in heat-treating the steel brace.

8. The method of claim 7, wherein the mid-section is cooled to 550° C. before removing the steel brace from the furnace used in heat-treating the steel brace.

9. The method of claim 8, wherein the heating and cooling treatments are completed within 2 hours.

10. The method of claim 1, wherein the mid-section region is heat-treated at a temperature of 700° C. for 30 minutes.

11. The method of claim 10, wherein the mid-section region is cooled at 0.6° C./min before removing the steel brace from a furnace used in heat-treating the steel brace

12. The method of claim 11, wherein the mid-section is cooled to 550° C. before removing the steel brace from the furnace used in heat-treating the steel brace.

13. The method of claim 1, further comprising cutting the steel brace to a final length, wherein the heat treatment of the steel brace is performed before the steel brace is cut to the final length.

14. The method of claim 1, further comprising cutting a hole or slot in the steel brace, wherein the heat-treatment of the steel brace is performed before the hole or slot is cut into the steel brace.

15. The method of claim 1, wherein an end of the mid-section region of the brace is at one brace depth away from an edge of the gusset plate.

16. A braced frame structure comprising: a heat-treated steel brace having a mid-section region of the steel brace with changed mechanical properties as compared to outer regions of the steel brace, the changed mechanical properties including improved material ductility, work hardening ability, and notch toughness; anda gusset plate coupled to the heat-treated steel brace without need of gusset plate reinforcement.

17. The braced frame structure of claim 16, further comprising a column or brace coupled to the gusset plate.

18. The braced frame structure of claim 16, wherein the steel brace comprises cold rolled steel.

19. The braced frame structure of claim 16, wherein the mid-section of the steel brace is relieved of residual stresses from prior cold work on the steel brace.

20. The braced frame structure of claim 16, wherein an end of the mid-section region of the brace is at one brace depth away from an edge of the gusset plate.