Large-Scale Testing of Low-Ductility, Concentrically-Braced Frames

C. Bradley1, J. Sizemore

2, J. Nelson

1, R. Tremblay

3, E. M. Hines

4, L. A. Fahnestock

5

1Graduate Research Assistant, Dept. Civ. & Env. Eng., Tufts University,

cbradley@lemessurier.com 2Graduate Research Assistant, Dept. Civ. & Env. Eng., UIUC

3Professor, Dept. Civ., Geo., & Mining Eng., École Polytechnique Montréal

4Professor of Practice, Dept. Civ. & Env. Eng., Tufts University;

Principal, LeMessurier Consultants 5Associate Professor, Dept. Civ. & Env. Eng., UIUC

ABSTRACT

In regions of moderate seismic hazard, the costly structural detailing necessary to

ensure adequate ductile performance of a braced frame during an earthquake event

can be difficult to justify economically. Structural engineers, however, are permitted

to design low-ductility systems if larger demands are assumed. In regions of moderate

seismic hazard, this philosophy has proved to be economical and has resulted in

widespread use of such low-ductility braced frame systems. Despite this popularity,

the inelastic behavior and collapse performance of these systems is not currently

understood at a fundamental level. A broadened understanding of the inelastic

behavior of low-ductility braced frames can lead to an improved seismic design

philosophy and provide practicing structural engineers with a coherent, rational, and

transparent design approach applicable to moderate seismic regions. This research

aims to identify low-ductility braced frame failure mechanisms and the sequence in

which they occur, as well as to draw conclusions on the implications of the observed

behavior contextually in building collapse performance.

INTRODUCTION

While steel concentrically braced frames constitute a significant portion of the

building stock in Eastern North America (ENA), these frames are generally designed

without consideration of seismic detailing with a response modification factor of R =

3, or as Ordinary Concentrically Braced Frames (OCBF) with a response

modification factor of R = 3.25. An OCBF design implies some consideration of

seismic detailing for the purposes of achieving more ductility than an R = 3 design

and less ductility than a Special Concentrically Braced Frame (SCBF), which has a

response modification factor of R = 6. Despite their prevalence, neither the R = 3 nor

the OCBF frame types have been tested to the levels of their more ductile

counterparts. Since 1997, when the AISC Seismic Provisions introduced the concept

of R = 3, the design criteria for OCBFs has changed in every update to the Seismic

Provisions (2002, 2005, 2010). Considering the fact that the OCBF system is one of

only two systems in ASCE 7 with an R-value written to three significant figures, and

considering that the value of 3.25 is only 8% higher than 3, the ambivalence of the

structural engineering community to the actual ductility of this system is quite clear.

In spite of the fact that so little is known about the seismic behavior and collapse

performance of low-ductility braced frames, the structural engineering community

has generally tolerated these frames in moderate seismic regions because of the

general resistance to collapse of steel braced frames during actual earthquakes, the

inherent ductility of structural steel as a material, the general level of redundancy in

most building systems with respect to collapse, and the understanding that buildings

are likely to form some kind of reserve system following the strength and stiffness

that results from brace fracture. The concept of reserve capacity has a long history in

seismic design, and was most clearly expressed in the original concepts behind dual

systems (Hines and Fahnestock 2010). Analysis of low-ductility braced frame

buildings with and without added reserve capacity in the form of column continuity

and beam-to-column moment-rotation capacity has also shown that enhanced reserve

capacity has a significantly greater effect on probabilistic collapse fragilities than

changes in strength of the braces and their connections (Hines et al. 2009).

Since this initial analytical work, however, several questions have been raised

regarding the sources of reserve capacity and their reliability. For instance, braces

that are assumed to reengage after initial fracture provide substantial reserve capacity

(Davaran et al. 2014, Sizemore et al. 2014). Brace gusset plate connections can

provide substantial beam-to-column moment capacity (Stoakes and Fahnestock

2011). Column continuity can also provide reliable reserve capacity when only braces

experience brittle failure in an isolated story.

The purpose of the large scale testing planned for the NEES facility at Lehigh

University this spring is to begin to provide some experimental data related to cyclic

behavior and brittle failure mechanisms of an R = 3 steel concentrically braced frame

and an OCBF. This paper outlines the goals of the testing program, the design of the

prototype, and the design of the test units considering realistic laboratory constraints.

NEES@LEHIGH TEST PROGRAM

During the spring and summer of 2014 two full scale tests of steel concentrically

braced frames (CBF) are underway at Lehigh University’s Advanced Technology for

Large Structural Systems (ATLSS) testing facility. This work is led by the University

of Illinois at Urbana-Champaign (UIUC) and Tufts University as part of a Network

for Earthquake Engineering Simulation Research (NEESR) project funded by the

National Science Foundation (NSF). UIUC and Tufts are collaborating with a team

from the École Polytechnique Montreal (EPM) on this work as part of an

international program to investigate reserve capacity in low-ductility concentrically

braced frames (Fahnestock et al. 2014). This program also includes testing and

analysis of top and seat angle connections for enhanced beam-to-column moment

capacity (Nelson et al. 2014, Beland et al. 2014), testing of tubular brace

reengagement with gusset plates after connection fracture (Davaran et al. 2014), and

collapse analysis of 3-, 6- and 9-story low-ductility braced frame buildings with

varying levels of reserve capacity (Sizemore et al. 2014). The NEES@Lehigh tests

include a two-story, R = 3, chevron brace configuration (Figure 1) and a two-story,

OCBF, split-X brace configuration (Figure 2).

Figure 1. R = 3 chevron braced frame test setup.

Figure 2. OCBF split-X braced frame test setup.

These tests are designed to explore post-elastic behavior in low-ductility braced

frames with a particular emphasis on brittle damage mechanisms. Over a year of

analytical work investigating the 3-story prototype buildings on which these two

frames are based led to the conclusion that while the overall purpose of the research

program is to investigate realistic collapse behavior of low-ductility braced frame

buildings under seismic loads, the purpose of these tests is to observe brittle damage

mechanisms for the purpose of calibrating analytical models. Considering the range

of behaviors that are possible in each frame and the scarcity of data on such frames,

the loading protocols and test setups for each frame are designed to be as simple as

possible.

The R = 3 chevron (or inverted-V) braced frame is generally considered to be an

archetypal low-ductility braced frame. It is widely used in building construction

throughout ENA because of the flexibility that it offers for architectural coordination.

It is also a statically determinate system that does not theoretically share gravity loads

between braces and columns. The AISC Seismic Provisions have always required

special consideration of such frames for OCBF designs such as: requiring the floor

beams to be able to carry gravity loads between columns without support from the

braces below, requiring the floor beams to resist not only gravity loads but also some

tension loads from braces after the compression brace has buckled, and requiring

stockier KL/r ratios for braces. While the Seismic Provisions have not considered

beam yielding as depicted in Figure 3(a) as a viable ductile mechanism for such

frames, such yielding has the potential to provide reserve capacity and will be the

primary focus of the R = 3 test. Floor beams in R = 3 systems are not required to span

between columns without help from the braces below. Nevertheless, the beams in this

frame have been designed both to span under gravity loads and to act as beam

columns under elastic lateral load transfer without consideration of bracing provided

by the floor slab.

Figure 3. NEES@Lehigh test unit post-elastic damage mechanisms.

(a) R = 3 chevron test unit, (b) OCBF split-X test unit.

Figure 3(a) and 3(b) illustrate the expected initial post-elastic behavior of the two test

units. In these figures, primary damage mechanisms are identified with larger circles

and secondary damage mechanisms are identified with smaller circles. The first two

damage mechanisms in both frames are expected to be buckling of the braces and/or

fracture of the compression brace connections. In the prototype building, compression

braces and connections are expected to fail first because their gravity loads and

seismic loads are additive. In the test unit, however, there are no gravity loads, so this

mechanism is not as obvious. Testing of tubular brace-to-gusset plate connections has

shown both that these connections can have actual strengths that exceed expected

strengths by more than 20%, and that minor accidental eccentricities can result in

significant bending of the gusset plates and tearing of the welds (Davaran et al. 2014).

For these reasons, brace buckling is expected prior to connection fracture, and the

investigators still expect to see post-elastic damage begin in the compression braces

and their connections. Inherent in damage mechanism 2 at the compression brace

connections is the possibility of brace reengagement after the weld has fractured and

the end of the slot bears against the edge of the gusset plate.

Damage mechanism 3 consists of yielding of the second floor beam once the

compression brace has buckled and this beam is subjected to an unbalanced load.

This damage mechanism is expected only in the R = 3 chevron frame. Damage

mechanisms 4 and 5 relate to the connections of the brace that is in tension during the

formation of damage mechanisms 1 and 2. Recent testing, conducted by Roeder et al.

of an older CBF typical of high seismic regions, demonstrated that this connection

fractured soon after the failure of the compression brace in buckling and low-cycle

fatigue fracture (Roeder et al. 2013). Of further interest as the Lehigh test units

degrade further in the post-elastic range are the performance of connections

associated with damage mechanisms 6 and 7. The bending in these mechanisms is

expected to be primarily a function of compatibility with overall frame deformations;

however, it may prove significant if it ultimately contributes to reserve capacity,

affects primary damage mechanisms, or experiences some kind of brittle failure that

further jeopardizes the integrity of the overall system. In particular, mechanisms 6

and 7 are expected to become active in the OCBF test unit if braces in both stories

fail. Such conditions would present an ironic twist on the split-X OCBF, rendering

the more intentional seismic system in greater need of reserve capacity in its beam-to-

column connections.

PRACTICAL LIMITATIONS OF THE NEES@LEHIGH TEST SETUP

Practical limitations of the testing equipment and the novelty of testing these brittle

systems have led to several decisions designed to simplify the test setups. The test

setups are full-scale so that member sizes, connection angles, gusset plates, bolts and

welds are inherently realistic. For the purposes of this discussion, the term “test

frame” refers to the Lehigh fixtures and stabilizing frame, and the term “test unit”

refers to a braced frame itself. For a full-scale test setup, the Lehigh test frame

offered a 35-ft bay, which is reasonably consistent with the 30-ft bay prototypes that

have been the subject of previous analysis. The available test frame story heights

were either 7.5 ft or 15 ft. Fifteen feet was chosen because of its general consistency

with designs for steel office buildings. This choice, consistent with the choice to test

at full scale, limited the test units to two stories and led to the decision not to include

column splices in the test units. While column splices had originally been conceived

as an important part of this testing, the investigators determined that such splices

would require backspans in the story both above and below the splice level,

necessitating a minimum of 3 stories. This would have either required the test units to

go to less than full scale, to not match any realistic prototype, or to significantly

change the stabilizing frames and test fixtures. None of these options were deemed to

be practical or necessary considering the number and complexity of behaviors already

expected in these test units.

Consistent with the Lehigh test frame, which stabilizes the test unit directly on either

side of the first and second floor beams, no concrete slabs were considered for the test

unit design. While concrete slabs provide significant post-brittle damage capacity in

steel buildings, this capacity typically depends on continuity of the slabs and decks

beyond the columns in question. Since the test unit consists of only one bay, the

added complexity of concrete slabs that would have to be coordinated with the test

frame and limited in width was not considered to be justified.

The test setup does not provide realistic gravity loads in the braced frame columns.

The test frames will experience a seismic-only load case. The lack of gravity loads

during the test will result in reduced stresses on the order of 15 ksi or 30% of yield

strength in the compression column. Although the stress variation in the columns is

significant, the lack of gravity in the columns is not expected to affect the braces and

connections as dramatically. Brace and connection failure are expected to occur

before column failure regardless of the presence of gravity loads. The absence of

gravity load combined with the likelihood and desirability of a first story mechanism

from an experimental point of view strengthened the argument for modeling the test

frames on a three story building structure as opposed to a taller structure. The three

story prototype was chosen instead of a two story prototype both for consistency with

previous analytical work and for the purpose of not including roof beams in the test

setup.

Columns in the test frames are connected via base plates to clevis fixtures, which

simulate ideal pins; however, the base reaction links intersect the test frame at an

elevation nearly 5 feet above the pinned base fixture. As a result, a pseudo-basement

level is formed below the first floor beam due to column continuity (Figure 4). The

resulting test unit column back spans are stiffer than they would be in a prototype

with a full basement. Analytical work to date has not modeled prototypes with

basements, opting to consider the first story as landing on spread footings with a slab

on grade. Neither realistic column base plates anchored into concrete footings, nor a

slab on grade were considered for the test unit design. Instead, at the ground level, a

steel grade beam was designed to transfer the base shear into an actuator with a load

cell that is braced to the laboratory floor.

Figure 4. Test unit base fixture and reaction link.

LOADING PROTOCOL

Each frame will be loaded quasi-statically with two actuators as shown in Figures 1

and 2. A third actuator is situated at the base beam to measure and resist the total base

shear. In the elastic range, the actuators will be force-controlled according to a regime

representing the equivalent lateral force distribution that served as the basis for the

test unit designs. In this range, the second floor actuator will be the master device,

loading the frame with both second and third story loads. The first floor actuator will

be slaved to the second floor actuator to maintain first floor loads that are

proportional according to the design force distribution. Since the exact moment or

location of failure is unknown, displacement limits will be set on both actuators that

are consistent with 0.2% total drift in the frame, which is equal to 0.72 in. at the

second story. If the system remains elastic up to this displacement level, the test will

proceed in displacement control in 0.1% total drift increments until the first damage

mechanism occurs. During this time, the lower actuator will remain slaved to the

upper actuator in linear proportion to the elastic force distribution.

Once the first damage mechanism occurs, the test will continue to proceed in

displacement control; however, the actuator loading the damaged story will become

the master actuator. For instance, if initial brittle damage commences in the first

story, the lower actuator will become the master. After initial brittle damage occurs,

the remainder of the test will be run in either 0.5% or 1.0% story drift increments for

the story containing damage. Loading of the undamaged story will be displacement-

based in proportion to the displacement recorded at the commencement of brittle

damage. If both stories experience brittle damage simultaneously, the investigators

will chose the master actuator based on the damage mechanism they prefer to

explore. In such a case, the slave actuator will run in linear proportion to its

displacement level recorded at the commencement of brittle damage.

DEVELOPMENT OF THE PROTOTYPE STRUCTURE

In a concerted effort to maintain a strong correlation between the somewhat artificial

test frames and actual braced frame designs, a prototype building was utilized in the

design process. The prototype structure was designed in accordance with ASCE-7

2005, the IBC 2006 and the Massachusetts state building code (8th

edition). Seismic

detailing for the OCBF followed the 2010 AISC Seismic Provisions instead of the

2005 Provisions, which are currently referenced by the Massachusetts Code.

The three-story prototype building (Figure 5) consists of a rectangular floor plan

(150ft x 175ft) with 15-ft story heights. There are two interior braced frames in each

direction, all utilizing strong-axis columns. The remaining columns are oriented such

that each half of the building contains an equal number of strong- and weak-axis

columns. Note that this configuration is intended to be as consistent as possible with

previous studies for steel frames in Boston (SAC 2000b, Hines et al. 2009).

Constraints in the ATLSS laboratory facility, however, required adjustments to bay

width and story height of the prototype structure. Specifically, the bay width was

increased from 30 to 35 feet in one direction, and all the story heights were adjusted

to 15 feet.

Figure 5 – Prototype building floor plan.

Composite slab construction of the prototype structure was assumed to be 3.25-inch

lightweight concrete slab on 3-inch steel deck. The total dead loads accounted for this

slab as well as deck, floor, hung, ponding, and framing loads. The total dead load was

taken to be 75 psf. Per ASCE 7, a 15 psf partition load was added to the baseline

office live load of 50 psf for a total live load of 65 psf. For the roof, dead and live

loads were each assumed to be 30 psf. Note that a façade load of 25 psf was assumed

on all exterior surfaces of the structure. In addition, a 10 psf live load for partitions

was added to the floor dead loads when determining seismic weights for calculation

of seismic loads.

Representative of the moderate seismic region in Boston, Massachusetts, the

prototype design assumes soil Site Class D and Seismic Design Category B in

developing the design acceleration response spectra. The seismic load effects were

determined by the equivalent lateral force procedure set forth in ASCE 7. Similarly,

wind loads were analyzed assuming the prototype to be located in Boston with

Expose Category B. The analytical procedure for wind loads, also presented in ASCE

7, was used for the design. Due to the relatively short height of the building, seismic

loads governed the design of the lateral system.

DEVELOPMENT OF THE TEST FRAMES

Two low-ductility lateral system test units were developed. One test unit will be a

two-story, full-scale, R=3, chevron brace configuration and the other a two-story,

full-scale, OCBF, split-X brace configuration. A summary of the framing members

and connection geometry are shown in Tables 1 and 2, respectively. Specified steel

grades are A992 for beams and columns, A500 Grade B for braces, and A36 for

angles and gusset plates. Additionally, all brace-gusset welds are 5/16 in. fillet welds.

Table 1. Test Frame Member Sizes and Beam Web Angle Sizes.

Table 2. Gusset Dimensions, Brace Weld Lengths, Beam Weld Sizes, and Gusset Angle Sizes.

Bracing connections in the frames consisted of square HSS profiles slotted around

and welded to gusset plates. The gusset plates, in turn, are welded to the connecting

beam. For connections at the supporting columns, the beams and gusset plates are

connected to the column flange by means of all-bolted double angles. In every double

angle connection, 3/4-inch diameter A490N bolts were specified. In order to evaluate

the distribution of brace forces to the connection elements, the Uniform Force

Method (UFM) was used. In addition, the double angles and bolts were analyzed for

prying action and designed to resist these effects.

Because the UFM provides little guidance on the final selection of gusset geometry, a

consistent approach was developed for use in the test unit designs. The final geometry

of the gusset plates can be determined by following these steps:

1. Determine the minimum required brace-to-gusset weld length necessary to

develop the ultimate brace design force. For this study, the weld capacity was

calculated assuming 1/4-inch fillet welds on all four corners of the brace. The

designs, however, specify 5/16-inch fillet welds to maintain single pass

welding while still accounting for a potential 1/8-inch of tolerance in the

width of the brace slot.

2. For connections at a support, assume a length for the gusset leg of the

Frame Story Beam Brace Web Angles Column

Chevron

R=3

3 W12x40 - 2L4x4x1/2

W1

2x

53

2 W12x40 HSS8x8x5/16 2L4x4x5/8

1 W12x72 HSS8x8x3/8 2L8x4x1 LLBB

OCBF

R=3.25

3 W12x40 - 2L4x4x1/2

W1

2x

53

2 W12x26 HSS6x6x3/8 2L4x4x1/4

1 W12x72 HSS6x6x1/2 2L8x4x1 LLBB

Fra

me

Sto

ry

Gu

sset

Po

siti

on

*

Gu

sset

Len

gth

, L

(in

)

Gu

sset

Hei

gh

t,

H (

in)

Gu

sset

Th

ick

nes

s, t

(in

)

Len

gth

of

Bra

ce-G

uss

et

Wel

d (

in)

Siz

e of

Bea

m-

Gu

sset

Wel

d

(in

)

Do

ub

le A

ng

le

Sec

tio

n

Chevron

R=3

3 B - S 16 14.5 3/8 6 1/4 2L4x4x3/8

2 B - S 19 17 1/2 9 5/16 2L4x4x1/2

T - M 48 16 1/2 9 3/8 -

1 B - S 21 19 1/2 11.5 5/16 2L4x4x1/2

T - M 51 17.5 5/8 11.5 7/16 -

OCBF

R=3.25

3 B - S 18 16.5 3/8 8 1/4 2L4x4x3/8

2 B - M 47 15.5 5/8 10.5 3/8 -

T - S 19 17.5 1/2 10.5 3/8 2L4x4x1/2

1 B - S 22 19.5 5/8 14.5 1/2 2L4x4x1/2

T - M 53 18 3/4 14.5 1/2 -

*Gusset positions are marked as Bottom/Top – Side/Middle and are abbreviated as B/T – S/M.

connecting angles. Although this length may change depending on prying

effects, 4 inches is reasonable starting point.

3. Assume a minimum desired spacing between the brace and the nearest

connecting element, which may be either the connecting angle or beam. The

purpose of this distance is to provide additional clearance for erection and

welding considerations.

4. Finally, assume an edge distance perpendicular to the brace that defines the

length of the cut in the gusset plate. In the test units, this value was selected to

be 2 inches because it resulted in reasonable final dimensions. Note that this

value cannot be less than the specified size of the brace-to-gusset fillet weld.

Assuming that the location of the work point of the brace and the angle of the brace

are known, then the above steps can be followed to define the gusset plate geometry

based solely on trigonometry. Once the gusset plate dimensions have been

determined, relevant limit states must be checked and the design adjusted as

necessary. Detail A from Figure 1 is the final design of the second floor support

connection in the chevron test unit and is displayed in Figure 6. This connection was

designed for ultimate brace forces equal to 191 kips in compression and 151 kips in

tension, which were determined from the 1.2D + 0.5L + 1.0E load combination.

General proportions of this connection are representative of typical braced frame

gusset plate connections.

Figure 6 – Chevron connection detail A.

The R=3, concentrically braced frame test unit is a low-ductility frame that considers

no seismic detailing. For this system, the members were designed for strength and

stiffness alone. Connections were not designed for the amplified seismic loads and no

amplification to the fundamental period, T, of the frame was considered.

The R=3.25 OCBF test unit is also a low ductility frame, but some detailing was

required. For V and inverted-V configurations, AISC 341 requires OCBF bracing to

satisfy b/t ≤ 15.4 and KL/r ≤ 96.3. In addition, connection elements, columns, and

beams must consider the amplified seismic load, which is directly proportional to Ω0.

For OCBF systems, ASCE 7 uses Ω0 = 2.0. Additionally, an increase to the

fundamental period was used in the analysis of seismic loads.

CONCLUSIONS

The results of this research will be applicable to braced frame design in moderate

seismic regions and may improve the work of design professionals in the field of

structural engineering as well as developers of seismic design and braced frame

design codes and specifications. The results will be of interest mainly to regions of

moderate seismic hazard, such as Eastern North America, but may also have

implications for design in high seismic regions. The results will provide valuable

insight into the behavior of these widely-used systems; thus, benefits to the structural

engineering profession will be significant.

ACKNOWLEDGMENTS

This study is supported by the National Science Foundation (Grant No. CMMI-

1207976) and the American Institute of Steel Construction. The opinions, findings,

and conclusions in this paper are those of the authors and do not necessarily reflect

the views of those acknowledged here.

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