BOLTED STEEL CONNECTIONS: TESTS ON T-STUB ... - CiteSeerX

BOLTED STEEL CONNECTIONS: TESTS

ON

T-STUB COMPONENTS

By James A. Swanson1 and Roberto T. Leon,2 Member, ASCE ABSTRACT: The results of tests on 48 T-stub specimens are presented and discussed. These tests were carried out as part of a SAC Phase II project in order to provide insight into the behavior, failure modes, and ductility of these components. The main variables tested include the size of the T-stub, the gauges of the bolts, and the type and diameter of the bolts. The primary intent was to develop a large database to calibrate simplified models suitable for design. In order to develop these models, a comprehensive instrumentation system that could identify the different components of deformation in the individual T-stubs was utilized. Most of the T-stubs failed by net section fracture through the stem and by tension fracture of the bolts, but generally the failure was after significant plastification had occurred. Other failure modes observed included bolt shear and block shear. The results indicate that current design equations provide conservative estimates of the ultimate strength of the Tstubs but that they are not necessarily good predictors of the governing failure modes.

INTRODUCTION

BEHAVIOR OF BOLTED CONNECTIONS

The numerous failures of fully welded moment connections during the 1994 Northridge and 1995 Kobe earthquakes have shown that conventional fully welded moment connections have several inherent drawbacks, which can be divided into two categories. The first category is related to the connection geometry, which promotes large strain demands in critical areas. The effects of these demands are often compounded by the sensitivity to fracture of typical welded connection details. Even if the welds are properly executed and made from tough materials, these strain demands will lead to tearing of the base metal in proximity of the weld access hole. This results in low rotational ductility and poor connection performance under large cyclic load reversals. The second category is related to the structural conception for modern steel moment frames, which requires, for economic reasons, that the lateral resistance be concentrated in relatively few connections. This results in heavy girders to control drift and in very large forces at the connections to maintain a weak beam–strong column mechanism. These large forces exacerbate the strain concentration and localization problems due to connection geometry. While the drawbacks related to connection geometry can be addressed through the use of tougher weld materials and better connection detailing, those related to structural conception remain. Thus, other connection types that avoid these problems should be investigated. Bolted T-stub connections are one alternative. Bolted and riveted connections have been used for decades and have performed well in past earthquakes, particularly when encased in concrete, as was traditionally done for fireproofing until the late 1950s (Munse 1976; Roeder et al. 1994, 1996; FEMA 1997). Using bolted connections throughout the structure provides a high level of redundancy and a level of stiffness comparable to that of fully welded connections. The research described in this paper attempts to develop criteria for the design and use of two types of bolted connections—T-stub and clip angle connections. Because of space limitations, the paper is limited to the discussion of the T-stub connection type.

Connection behavior can be conveniently represented by a moment-rotation curve, as shown in Fig. 1. Connections can be classified by three main characteristics: strength, stiffness, and ductility. As far as strength is concerned, connections are classified as either full strength (FS) or partial strength (PS) depending on whether they are capable of transferring the full plastic moment of the framing beams. As far as stiffness is concerned, connections are classified as full restraint (FR), partial restraint (PR), or simple connections depending on the service load stiffness. Finally, connections are classified as brittle or ductile and certified for use in ordinary, intermediate, and special moment frames based on their ability to reach and sustain certain plastic rotational demands. In the aftermath of the Northridge earthquake, a plastic rotation of 0.03 rad under cyclic loading, with a loss of strength less than 20%, has been suggested as an acceptable rotation limit to differentiate between ductile and brittle connections for special moment-resisting frames. The primary goal of this project was to develop design rules for T-stub connections that would result in a full-strength connection, ductile behavior, and a connection stiffness close to but not necessarily in the FR range. This type of behavior for a T-stub connection was deemed reasonable if the beam flange size was kept low (W18 to W27, with weights less than 100 lb/ft). Although T-stubs have been used successfully for many years, most of the research has focused on monotonic loading and issues related to prying action [see Batho and Rowan (1934); Douty and McGuire (1965); and Agerskov (1976, 1977), for example]. A summary of bolted connection research for seismic applications can be found in FEMA (Leon 1997). Previous research on T-stub behavior under cyclic loading is

1 Asst. Prof., Dept. of Civ. and Envir. Engrg., Univ. of Cincinnati, Cincinnati, OH 45221-0071. 2 Prof., School of Civ. and Envir. Engrg., Georgia Inst. of Technology, Atlanta, GA 30332-0355. Note. Associate Editor: Sashi Kunnath. Discussion open until June 1, 2000. To extend the closing date one month, a written request must be filed with the ASCE Manager of Journals. The manuscript for this paper was submitted for review and possible publication on April 5, 1999. This paper is part of the Journal of Structural Engineering, Vol. 126, No. 1, January, 2000. qASCE, ISSN 0733-9445/00/0001-0050–0056/$8.00 1 $.50 per page. Paper No. 20627.

50 / JOURNAL OF STRUCTURAL ENGINEERING / JANUARY 2000

FIG. 1.

Typical Moment-Rotation Curve

scant. The most recent and relevant research in this area is that of Roeder et al. (1994, 1996) and Bursi et al. (1996). The results of the studies by Roeder et al. indicated that the behavior of the T-stub connections was strongly influenced by local slipping of the bolts, gauge distances, yielding of the unreinforced panel zone, and concrete encasement when present. CONNECTION MODELING The complex nature of bolted connection behavior, as evidenced by the nonlinear nature of most moment-rotation curves shown in Fig. 1, requires that advanced models be used if proper account of the effects of connection behavior is to

FIG. 2.

be taken in design. These advanced models require either continuous or multilinear moment-rotation curves for each connection. There are currently three methods for obtaining these curves: experimentation, advanced finite-element modeling, and curve fitting to existing test data. None of these methods by itself provides a safe and efficient solution. A fourth alternative for obtaining a moment-rotation curve is the component spring model initially used by the Eurocode (1992). In this approach, the individual components of the connection are modeled by linear or nonlinear springs. Each of these springs is then added to the system and its stiffness is assembled into the final, overall rotational stiffness of the connection. Fig. 2 shows a spring model for a complete connection, and Fig. 3 shows a spring model of an individual T-stub. The overall connection stiffness typically consists of the T-stub stiffness, web angle or web plate stiffness, panel zone stiffness, and the connection region of the beam. The model of the individual T-stub typically consists of the contribution from (1) the tension bolts; (2) the T-stub flange; (3) the T-stub stem; (4) the shear bolts; (5) the bearing deformations; and (6) the connection slip. The different stiffnesses in the T-stub model may be combined in parallel or series depending on how they interact with each other. This approach offers the possibility of combining experimental results and mechanistic models to arrive at a relatively simplified model for connection design, TABLE 1.

Spring Model of T-Stub Connection

Material Properties

Series (1)

Mill Fy (2)

Mill Fu (3)

TA TB TC TD CA

365.4 MPa 379.2 MPa 406.8 MPa unknown 379.2 MPa

482.6 MPa 479.2 MPa 513.7 MPa unknown 527.4 MPa

Coupon Fya (4) 352.3 378.5 415.1 426.8 360.3

MPa MPa MPa MPa MPa

Coupon Fua (5) 468.8 500.6 530.2 569.5 528.8

MPa MPa MPa MPa MPa

a

FIG. 3.

Spring Model of Individual T-Stub

FIG. 4.

Coupons taken from web of material.

Typical T-Stub Dimensions JOURNAL OF STRUCTURAL ENGINEERING / JANUARY 2000 / 51

and was selected for use in this study. Thus, the experimental program described next was intended to provide the necessary data to develop a robust component spring model for a T-stub connection. TABLE 2.

EXPERIMENTAL PROGRAM As a first step in this research, a comprehensive literature review was compiled. The database included 771 bolted con-

Component Test Descriptions


nection tests, of which 11% were T-stub tests and 19% were conducted using cyclic load histories. The database was used primarily to identify gaps in the experimental data and to plan the component test series. The final experimental program consisted of 48 T-stubs and 10 clip angles to be tested individually under cyclic and monotonic loads. Additionally, six full-scale beam-column tests were conducted and used to calibrate the overall connection model. Only the T-stub tests are addressed in this paper. Information on the other tests, as well as substantial additional information on the T-stub tests, can be found at the web site http://www.ce.gatech.edu/;sac/. The experimental program is centered around testing individual T-stubs, such as those shown in Fig. 4, rather than full connections. By testing individual T-stubs or components instead of entire beam-column subassemblages, a large number of parameters was investigated economically and rapidly. The components were subjected to axial loads based on expected beam flange forces in actual connections. The axial loading is a simplification of the actual conditions, since both localized bending and shear forces present in the actual connection were missing in the component tests. The T-stubs were cut from four different wide flange sections and were designed to be used in full strength connections for smaller W18 to W27 beam sections. However, they would also be suitable as partial strength connections for larger W30 to W36 beam sections. Based on current code-type calculations, bolt shear strength often governed the trial designs of the T-stubs, since it would be difficult to place more than eight bolts in the stem of a typical WT. As a result, the T-stubs used in this research were cut from W shapes, in a castellated fashion, to allow most of the web of the section to be used as a stem for the T-stub. This allowed more shear bolts to be placed in the extended stem and reduced the number of cases where bolt shear controlled the design. All of the components were fabricated from A572 Grade 50 steel. The material properties for the components are shown in Table 1 and the section properties are provided in Table 2. The T-stubs were designed for a typical column and were thus cut 394 mm (15.5 in.) wide at the tension flange. The stems of the T-stubs were tapered to the width of the beam flange. Fig. 4 shows the dimensions of a typical T-stub. The size, grade, number, gauge, and spacing of the bolts were varied to study the effects of prying on the tension flange and bearing on the stem. Various configurations of 22 and 25 mm (7/8 in. and 1 in.) diameter A325 and A490 tension control bolts were used. Series A, B, and D T-stubs were designed for a beam set back of 51 mm (2 in.), and series C T-stubs were designed for a set back of 57 mm (2.25 in.). Table 2 provides a complete description of the component test specimen and a summary of results. Most of the specimens were subjected to cyclic load histories applied axially. Four duplicate T-stubs were tested monotonically to provide reference points for comparing the cyclic data to the existing monotonic data found in the literature search. The cyclic load history specified in the SAC (1997) testing protocol was used. It consisted of several steps of increasing rotations or displacements, each made up of several cycles. Because of the different stiffnesses of the components in tension and compression, a system of force and displacement limits was used for the cyclic tests. During the tension portion of the first cycle of a given step, the T-stub was pulled to a given displacement and the load was recorded. The specimen was then pushed in compression until a load equal to the tensile load was reached. This system of force and displacement limits, while in stroke control, was used for all of the cyclic tests. Component testing was conducted using a specially constructed frame. Four actuators were mounted in parallel between two spreader beams. This set-up allowed com-

FIG. 5.

Instrumentation Scheme for Component Tests

ponent loads of up to 3,100 kN (700 kips) to be applied. The T-stubs and clip angles were bolted to a stiffened W14 3 257 column stub that was in turn bolted to the lower spreader beam. The T-stem or outstanding angle leg was then bolted to a longer T-section, spanning between the spreader beams, that was used to represent the beam. Fig. 5 shown the instrumentation employed during the component tests. Linear variable displacement transformers (LVDTs) and potentiometers were used to isolate the different components of the overall deformations. LVDT A monitored the connection slip, B measured uplift of the T-flange from the face of the column, C measured uplift of the T-flange at the bolt line, D measured elongation of the T-stem, E measured the overall T-stub deformation, and G indicates the use of bolts instrumented with strain gauges to monitor bolt forces, including the effects of prying. Strain gauges were also used on the T-stubs to monitor strains. BEHAVIOR Fig. 6 shows a force-deformation curve typical of T-stubs tested cyclically. Notice that the deformations are larger in tension than in compression as a result of the testing procedure discussed previously. A connection rotation can be estimated

FIG. 6.

Cyclic Force-Deformation Curve

JOURNAL OF STRUCTURAL ENGINEERING / JANUARY 2000 / 53

by dividing the axial displacement range by the expected beam depth to obtain an angle. In this case, a rotation of 0.035 rad can be approximated for a W24 beam from the peak to peak displacement of about 22 mm (0.84 in.). Fig. 7 shows the slip behavior of a T-stub tested cyclically. Notice that the slip resistance deteriorates throughout the test and that the entire curve lies to the left of center. This shift is due in part to the alignment of the bolt holes of the beam flange and T-stem before testing. Because of this, not an absolute value but the deformation range of a T-stub must be used to evaluate performance. Fig. 8 shows the stem deformation of a T-stub tested

FIG. 10.

FIG. 7.

Cyclic Force-Slip Curve

Monotonic versus Cyclic Behavior

cyclically. The overall plastic deformation of the stem is tensile and shows a limited amount of hysteresis. Fig. 9 shows the flange uplift of a typical T-stub tested cyclically. The deformations of this T-stub (TA-03) were well balanced between the stem and flange. The major energy dissipating mechanisms were the flange uplift and connection slip. Fig. 10 shows the force-deformation of a T-stub (TA-07) tested monotonically superimposed on the force-deformation curve from Fig. 6. The two T-stubs were identical except for the load history. Superimposing the two curves shows that the monotonic data provides an accurate envelope of the cyclic data. Comparisons of the summations obtained for most of the T-stubs to the measured total deformations indicated that in most cases the instrumentation worked well. TENSION BOLT AND FLANGE STRENGTH

FIG. 8.

Cyclic Force-Stem-Deformation Curve

FIG. 9.

Cyclic Force-Flange Uplift Curve


Twenty-one T-stubs failed in a tension bolt fracture mode. The tension bolt failures were the most sudden and brittle of the failure modes observed. The only warning of failure was prying or uplift on the flange of the component. Fig. 11 shows the behavior of four T-stubs with tension bolt gauges ranging from 101 mm (4 in.) to 178 mm (7 in.), in increments of 25 mm (1 in.). TA-01, with a tension bolt gauge of 101.6 mm (4 in.), was the stiffest of the four. The stiffness of the flange shifted the failure of the T-stub into the stem, as opposed to others that experienced tension bolt failures. TA-04 was the most flexible of the group, with a tension bolt gauge of 177.8 mm (7 in.), and failed in bolt tension fracture after developing a flange mechanism (prying action).

FIG. 11.

Flange Uplift Behavior

FIG. 12.

LRFD Prying Model Accuracy

Several prying models for predicting the ultimate strength of T-stub connections are available. Most are based on work by Douty and McGuire (1965), Nair et al. (1974), Agerskov (1976, 1977), Thorton (1985), or Kulak et al. (1987). Fig. 12 shows a comparison of the strength obtained experimentally to that predicted by the current LRFD approach for these 21 tests. The LRFD approach is based on the work of Fisher et al. (1987) and Thornton (1985), and the values used for the comparison were calculated using a resistance factor of 1.0 and the actual material properties. The figure shows the ratio of (predicted strength–actual strength) to the actual strength. A negative error is conservative, and in all cases the predicted capacity was lower than the actual observed capacity. The average error was 27.7%.

FIG. 13.

Whitmore Width

STEM STRENGTH Twenty-four T-stubs failed with net section fractures of the stem, and one T-stub developed a block shear failure. Because the column flanges were wider than the flanges of the beams, the stems of all of the T-stubs were tapered. Depending on the severity of this taper, the entire width of the stem may or may not be used in the capacity calculations. An angle of 307 measured from the first row of bolts was used to define the effective width or Whitmore width, as shown in Fig. 13 (LRFD 1994, Volume II, pp. 11–15). The smaller of this width and the actual width was then used to calculate the net area used in fracture strength calculations. The accuracy of the model is shown in Fig. 14. The data is shown in the same format as for Fig. 12, with a negative error being conservative. The predicted capacities were all higher than the actual values. T-stub TA-26 failed in a block shear mode, as shown in Fig. 15. In the design of the test series, shear bolt fracture often governed the strength of the T-stubs. As a result, T-stubs TA09, TA-25, and TA-26 were designed to investigate the possibility of reducing the shear bolt spacing, enabling more bolts to be placed in the stem. The spacing of the shear bolts on TA-09 was 3db, as is recommended in the LRFD manual. The shear bolt spacing of T-stub TA-25 was reduced from 3db to 2 2/3db, as is permitted by LRFD. The spacing on TA-26 was reduced from 3db to 2 1/2db. This reduced bolt spacing had little effect on T-stubs TA-09 and TA-25 but resulted in a block shear failure for TA-26. The mechanism of block shear consists of either (1) shear yielding accompanied by tensile fracture or (2) shear fracture accompanied by tensile yielding. Using equation J4-3 of the current LRFD, the capacity of TA-26 is 1,820 kN (409 kips) for case (1) or 1,522 kN (342 kips) for case (2). The failure started when a crack formed in the stem between the last two bolts, as in the net section fractures, and then shear yielding along the bolt lines progressed until the tear out was complete.

FIG. 14.

Net Section Fracture Model Accuracy

FIG. 15.

Block Shear Failure

Based on these observations, case (1) should govern. The LRFD, however, defines the governing case as that with the larger fracture term, or case (2). The actual failure load for TA-26 was 1,775.7 kN (400 kips). JOURNAL OF STRUCTURAL ENGINEERING / JANUARY 2000 / 55

TABLE 3.

Shear Bolt Failures

Test ID (1)

Predicted (2)

Actual (3)

TB-01 TB-03

2249.0 kN 1866.9 kN

2252.1 kN 2108.0 kN

tent reliability factor, cannot be used to rank the failure modes at ultimate. • The capacity of the T-stem can be increased by using a larger spacing between the rows of shear bolts. This reduces the taper of the stem and increases the effective area or Whitmore section.

SHEAR BOLT STRENGTH

ACKNOWLEDGMENTS

Two T-stubs failed when the shear bolts fractured. All of the T-stubs were designed with the shear bolts in the ‘‘threads excluded’’ condition to avoid shear bolt failures. In this case, the capacity of a shear bolt is calculated as the product of Ab and Fv, as shown in section J3.6 of the current LRFD. The actual and predicted capacities are compared in Table 3. The predicted capacity closely matches that observed for test TB01, and the predicted value is somewhat conservative when compared to the results of test TB-03.

The work described in this paper is being sponsored by the Federal Emergency Management Agency (FEMA) through a grant to SAC. The material donations by NUCOR Corporation, LeJeune Bolt Co., and Cives Steel Co. are gratefully acknowledged.

BEAM-COLUMN TESTS Six out of a total of 10 full-scale beam-column tests have also been completed. These tests were designed as benchmarks for relating the component test data to useful connection data. Preliminary comparisons indicate that the T-stubs in the fullscale tests behaved very similarly to those tested as components. However, extensive local buckling of the beam flange and web governed the behavior in most tests. The results of these tests, which will be the subject of a follow-up paper, are displayed at the web site http://www.ce.gatech.edu/;sac/. CONCLUSIONS Based on the results of the 48 T-stub tests described in this paper, the following conclusions can be made: • The major contributions to the overall deformation of Tstub components were made by the flange deformation, tension bolt elongation, stem deformation, and relative slip. These could be measured independently and correlate well with simplified models. • The major energy dissipating mechanisms were the flange deformation (yield) and connection slip. • The stiffness of the T-stub flange can be greatly increased by decreasing the tension bolt gauge or by increasing the flange thickness. • Existing strength models in the LRFD provide adequate predictions for the strength of T-stub components. The predicted flange and tension bolt capacities were shown to be slightly conservative, while the net section fracture and shear bolt fracture capacities were shown to be slightly unconservative. • Current design equations, which do not contain a consis-


APPENDIX.

REFERENCES

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